BASIC TESTING AND STRENGTH DESIGN OF CORRUGATED BOARD AND CONTAINER CONTAINERS S
TOMAS NORDSTRAND
Structural Mechanics
Doctoral Thesis
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Structural Mechanics ISRN LUTVDG/TVSM--03/1015--SE (1-129) ISBN 91-628-5563-8 ISSN 0281-6679
BASIC TESTING AND STRENGTH DESIGN OF CORRUGATED BOARD AND CONTAINERS
Doctoral Thesis by TOMAS NORDSTRAND
Copyright © Tomas Nordstrand, 2003. Printed by KFS i Lund AB, Lund, Sweden, February 2003. For information, address: Division of Structural Mechanic s, LTH, LTH, Lund University, Box 118, SE-221 00 Lund, Sweden. Homepage: http://www.byggmek.lth.se
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ABSTRACT
Packaging serves a lot of purposes, and would be hard to do without. Packaging protects the goods during transport, saves costs, informs about the product, and extends its durability. A transport package is required to be strong and lightweight in order to be cost effective. Furthermore, it should be recycled because of environmental and economical concerns. Corrugated board has all of these features. This thesis is compiled of seven papers that theoretically and experimentally treat the structural properties and behaviour of corrugated board and containers during buckling and collapse. The aim was to create a practical tool for strength analysis of boxes that can be used by corrugated board box designers. This tool is based on finite element analysis. The first studies concerned testing and analysis of corrugated board in three-pointbending and evaluation of the bending stiffness and the transverse shear stiffness. The transverse shear stiffness was also measured using a block shear test. It was shown that evaluated bending stiffness agrees with theoretically predicted values. However, evaluation of transverse shear stiffness showed significantly lower values than the predicted values. The predicted values were based on material testing of constituent liners and fluting prior to corrugation. Earlier studies have shown that the fluting sustains considerable damage at its troughs and crests in the corrugation process and this is probably a major contributing factor to the discrepancy. Furthermore, the block shear method seems to constrain the deformation of the board and consistently produces higher values of the transverse shear stiffness than the three-point-bending test. It is recommended to use the latter method. Further experimental studies involved the construction of rigs for testing corrugated board panels under compression and cylinders under combined stresses. The panel test rig, furnishing simply supported boundary conditions on all edges, was used to study the buckling behaviour of corrugated board. Post-buckling analysis of an orthotropic plate with initial imperfection predicted failure loads that exceed the experimental values by only 6-7 % using the Tsai-Wu failure criterion. It was confirmed, by testing the cylinders that failure of biaxially loaded corrugated board is not significantly affected by local buckling and that the Tsai-Wu failure criterion is appropriate to use. A method for prediction of the top-to-bottom compression strength of corrugated board containers using finite element analysis was developed and verified by a large number of box compression tests. Up to triple-wall corrugated board is accommodated in the finite element model. The described FE-method for predicting the top-to-bottom compressive strength of corrugated containers has been used as the basic component in the subsequent development of a user-friendly computer-based tool for strength design of containers. Keywords: analysis, bending, box, buckling, collapse, compression, corrugation, corrugated board, crease, design, experiment, failure criterion, fluting, finite element method, liner, local buckling, packaging, panel, paper, stiffness, strength, test method, transverse shear
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CONTENTS PART I: INTRODUCTION AND SUMMARY
PAGE 1
General remarks
1
Background and earlier work
3
Aim of present work
4
General assumptions and limitations in present work
5
Summary of contents and major conclusions
5
Concluding remarks and future research
6
Presented papers
7
Acknowledgements
8
References
9
PART II: APPENDED PAPERS
PAPER 1: T. Nordstrand, H. G. Allen and L. A. Carlsson, "Transverse Shear Stiffness of Structural Core Sandwich", Composite Structures, No. 27, pp. 317-329, 1994. PAPER 2: T. Nordstrand and L.A. Carlsson, "Evaluation of Transverse Shear Stiffness of Structural Core Sandwich Plates", Composite Structures, Vol. 37, pp. 145-153, 1997. PAPER 3: T. Nordstrand, "On Buckling Loads for Edge-Loaded Orthotropic Plates including Transverse Shear", SCA Research, Box 716, 851 21 Sundsvall, Sweden. To be submitted to Composite Structures. PAPER 4: T. Nordstrand, "Parametrical Study of the Post-buckling Strength of Structural Core Sandwich Panels", Composite Structures, No. 30, pp. 441451, 1995. PAPER 5: T. Nordstrand, "Analysis and Testing of Corrugated Board Panels into the Post-buckling Regime", SCA Research, Box 716, 851 21 Sundsvall, Sweden. To be submitted to Composite Structures. PAPER 6: P. Patel, T. Nordstrand and L. A. Carlsson, "Local buckling and collapse of corrugated board under biaxial stress", Composite Structures, Vol. 39, No. 1-2, pp. 93-110, 1997. PAPER 7: T. Nordstrand, M. Blackenfeldt and M. Renman, "A Strength Prediction Method for Corrugated Board Containers", Report TVSM-3065, Div. of Structural Mechanics, Lund University, Sweden, 2003.
Part I Introduction and Summary
INTRODUCTION AND SUMMARY General remarks
In 2001 the European transport packaging market had an estimated value of approximately $20 billion. Corrugated board represented 62 p er cent of this market value [1]. A transport package is required to be strong and lightweight in order to be cost effective. Furthermore, it should be recycled because of environmental and economical concerns. Corrugated board has all of these features. In its most common form, viz. single-wall board, two face sheets, called liners, are bonded to a wave shaped web called fluting or medium, see Figure 1. The resulting pipes make the board extremely stiff in bending and stable against buckling in relation to its weight [1]. Consequently, the strength of the wood fibres in the board is also utilised in an efficient way. The fluting pipes are oriented in the cross-direction (y, CD) of board production, see Figure 1. The orientation of the board in-line with production is called machine-direction (x, MD). Orientation through the thickness of the board is denoted Z-direction (z, ZD). This definition of principal directions is also used for the constituent paper sheets. z, ZD y, CD X, MD
Figure 1. Single-wall corrugated board. In area, about 80 per cent of corrugated board production is single-wall board. The rest is produced for more demanding packaging solutions that require double or triple-wall board, illustrated in Figure 2. z, ZD y, CD X, MD
Figure 2. Double and triple-wall corrugated board. The profile of a corrugated web in Figure 3 is characterised by a letter, A, B, C, E or F, specified in Table 1 [1]. Also listed in Table 1 are the take-up factors which quantify the length of the fluting per unit length of the board. For example, one metre of corrugated board with B-flute requires a 1.32 m long piece of paper prior to corrugation.
1
hc
z, ZD y, CD X, MD
λ
λ
Figure 3. The geometry of a corrugated web. As seen in Table 1 the tallest core profile is A-flute, which is used in board for heavy duty boxes. B and C-flute are used for the most common board grades. The E and Fflutes are small and consequently used in board for smaller boxes, e.g. perfume packages, where appearance and printability are important [1]. Table 1. Flute profiles. Profile Wavelength, λ (mm) Flute height, h c (mm) Take-up factor, α
A 8.3-10 4.67 1.54
B 6.1-6.9 2.46 1.32
C 7.1-8.3 3.61 1.43
E 3.2-3.6 1.15 1.27
F 2.3-2.5 0.76 1.25
A corrugator is a set of machines in line, designed to bring together liner and medium to form single, double or triple-wall board. This operation is achieved in a continuous process, see Figure 4. The reels of liner and medium are fed into the corrugator. The medium is conditioned with heat and steam and fed between large corrugating rolls forming fluting. In the Single Facer, starch adhesive is applied to the tips of the flutes on one side and the inner liner is glued to the fluting. The fluting with one liner attached to it is called single-face web and travels along the machine towards the Double Backer where the single-face web is bonded to the outer liner and forms corrugated board. The corrugated board is then cut and stacked. Double Backer Corrugated board
Single Facer Liner
Machine direction MD
Fluting
Liner
Medium
Figure 4. Manufacture of corrugated board.
2
The first corrugators were built in the US at the start of the last century. However, up until 1920, the majority of products shipped via railroads, for example, were packed in wooden crates. The corrugated box was relatively new and few had any experience in transporting them. In order to avoid liability for damage while shipping items in corrugated boxes, railroads in the US established a standard known as Rule 41. Rule 41 was an important step in opening up the market for corrugated board packaging. Later on, during World War II, corrugated board packaging was called upon to deliver rations and other war material to all corners of the earth. This contributed to the establishment of corrugated board globally. After the war the market grew rapidly, and the range of sizes and capabilities of corrugated boxes grew to fit the myriad of new products developed. Recently, the combination of a plastic bag inside a corrugated board box (bag-in-box) has resulted in many new opportunities, including the latest trend packaging of wine. Corrugated board is permeable to moisture and absorbs water. This will reduce its strength and stiffness. However, it can be made both water and grease proof. Many package styles and design options are possible, but often an international standard of box styles [2], the FEFCO-code, is used in specifying a design. One of the most common box styles is the regular slotted container (RSC) denoted FEFCO 0201, see Figure 5. The box size is specified by L xWxH, i.e. length of the longest side panel, width of the shortest side and height. The flap size is half of the width. In the logistics chain in Sweden a transport package is usually adjusted to the EUR-pallet. Thus the length and width of an RSC are usually uniform divisions of the pallet size (1200x800 mm), e.g. 300x200 mm or 600x400 mm. Crease Flap W/2 H L
W
L
W
H
Side panel
W/2 L W
Figure 5. A regular slotted container, code FEFCO 0201. RSC:s are produced with an in-line Slotter-Folder-Gluer, which in one operation creases, cuts, folds and glues the blank into its final shape. The RSC is then palletised and ready to be shipped flat to the customer .
Background and earlier work
Several experimental studies have been conducted on the compression strength of corrugated board containers [3,4]. The most common failure mode for a corrugated box loaded in top-to-bottom compression is post-buckling deflection of its side panels,
3
followed by biaxial compressive failure of the board in the highly stressed corner regions of the box. Local instabilities of the liners and fluting may also interact with the failure progression [5-8]. A detailed finite element analysis of a corrugated board panel has shown that local buckling of one of the liners may occur before actual material failure [9]. This can also be observed visually just prior to compression failure of panels and boxes [10]. However, for shallow boxes and boxes with high board bending stiffness in comparison to the box perimeter, failure is often caused by crushing of the creased board at the loaded edges instead of collapse during buckling [11]. When considering the compression of panels in a box it is recognised that the flaps, attached to the panels through the creases at top and bottom edges, introduce an eccentricity in the loading [12, 13]. Furthermore, the top and bottom edges normally have a much lower stiffness than the interior of the panel due to the creases. It has been concluded that the low stiffness prevents a redistribution of the stresses to the corners of the box and consequently reduces the box compression strength. Several previous investigations have involved finite element analysis of corrugated board. Peterson [14] developed a finite element model to study the stress fields developed in a corrugated board beam under three point loading. Pommier and Poustius studied bending stiffnesses of corrugated board using a linear elastic finite element code [15]. Pommier and Poustius also developed a linear elastic finite element model for prediction of compression strength of boxes [16]. Likewise a linear elastic finite element model of a corrugated board panel for prediction of compression strength was developed by Rahman [17]. Patel developed a linear elastic finite element model in a study of biaxial failure of corrugated board [18]. The model was used to predict buckling patterns of a circular tube subjected to different loading conditions. In an investigation by Nyman, local buckling of corrugated board facings was studied numerically through finite element calculations [19]. Little published work is available on the use of non-linear constitutive models for prediction of strength of corrugated board structures. However, a non-linear model of corrugated board was developed by Gilchrist, Suhling and Urbanik [20]. In their model, both material and geometrical non-linearities were included, in-plane and transverse loadings of corrugated board were examined. Bronkhorst and Riedemann [21] and Nordstrand and Hagglund [22] have developed non-linear finite element models for corrugated board configurations. These investigations generated predictions for compressive creep of a box and time-dependent sagging of a corrugated board tray.
Aim of present work
This project was initiated with the objective of developing a design method based on fundamental engineering mechanics to predict the strength of corrugated containers in top-to–bottom compression.
4
General assumptions and limitations in present work
The major assumptions and limitations adopted in this work are as follows: • • • •
Paper is regarded as a homogenous continuum with liner elastic orthotropic properties. Influence of load duration, e.g. creep, moisture and inertia forces are not analysed. Deterministic characterisation of material properties, geometry and loading. Box strength analysed only for loading in top-to-bottom compression.
Summary of contents and major conclusions
In Paper 1, expressions for the transverse shear stiffnesses of corrugated board are derived by considering a shear loaded element of the corrugated board and using the theory of curved beams. It is shown how the transverse shear stiffness in the machine direction is significantly changed by the transition from one core shape to another. An experimental study of the transverse shear stiffness is given in Paper 2, where the transverse shear stiffness is measured both by a block shear test and evaluated from a three-point flexure test. The three-point flexure test is also simulated using finite element analysis. Values of transverse shear stiffnesses obtained from the block shear test are much larger than values evaluated from the three-point flexure test. The difference is attributed to the highly constrained deformation of the facings in the block shear test. It is also shown that experimental values are significantly lower than calculated values obtained in Paper 1 and obtained from the finite element analysis. This is probably caused by delamination damage to the corrugated medium inflicted during the corrugation process. In Paper 3, an expression is derived for the buckling load of a simply supported orthotropic plate including first order transverse shear deformation. The influence of the transverse shear on critical buckling is studied and compared with ordinary sandwich theory. Its primary use, however, is to verify the buckling load obtained in a finite element analysis of a simply supported single-wall corrugated board panel in Paper 4. The influence on the panel strength of different parameters such as asymmetry, slenderness of the corrugated board and eccentric loading is studied in Paper 4. It was concluded that panel strength is very sensitive to boundary conditions and change in core thickness of the board, i.e. the change in bending stiffness of the board. In Paper 5, a panel compression test rig, furnishing simply supported boundary conditions on all edges, was designed and used to study the buckling behaviour of corrugated board panels. An analysis of an orthotropic plate with initial imperfection is presented in Paper 5 to predict the collapse load using the Tsai-Wu failure criterion. A significant difference was observed between analytically predicted and experimentally measured displacements at large out-of-plane deformation. This is probably caused by non-linear material behaviour of paper and local buckling of the panel facings, i.e. the 5
liners. However, the analytically predicted failure load exceeds experimental values by only 6-7 %. This suggests that collapse of the corrugated board panel is triggered by material failure of the inner facing. It is also concluded in Paper 6, where an experimental study of biaxially loaded corrugated board is presented, that failure is not significantly affected by local buckling of the corrugated board and that the Tsai-Wu criterion is appropriate to use. Finally, in Paper 7, a finite element method developed for stress and strength analysis of corrugated containers using the failure criterion above is presented. The corrugated board is represented by multi-ply eight node isoparametric shell elements, and the soft creases at the loaded top and bottom edges are accommodated in the finite element model by spring elements. Effective material properties of the homogenised corrugated cores have been used, and each layer of the corrugated board is assumed to be orthotropic linear elastic. It is shown that convergence is obtained with relatively few elements, e.g. 144 elements are quite sufficient for a regular size box, i.e. 300 x300x300 mm. Sensitivity of the collapse load to the imposed compliance at the loaded boundaries is also studied. Different buckling modes of a box are simulated giving an in-depth understanding of the relation between the strength of a box and constraints imposed on the panels by the corners of the box. Extensive testing of boxes made from B- and Cboard shows that predicted failure loads using the proposed finite element model have an average error margin of 5% compared to measured box strengths.
Concluding remarks and future research
Box performance requirements range from its appearance, to its mechanical strength and ability to protect its contents. Mechanical properties can be divided into two categories, those that pertain to rough handling and stacking. Both of these types are difficult to duplicate accurately in the laboratory. As a consequence, the box compression test or BCT of an empty container has been widely used as a means of evaluating container performance. However, in order to distinguish between factors that govern box performance it is necessary to test the quality of the corrugated board and its components, maintain good control of conversion operations and environmental influences such as humidity and load duration. In addition to standard testing methods, a future challenge for research is to develop more sophisticated testing methods that are based on finite element models. Once the roles of liner and medium behaviour in box performance are properly understood, material properties can be evaluated by mill and plant personnel so that attention is given to the properties that govern end-use performance. For example, corrugated containers that are stacked on top of each other will slowly deform with time until one of the boxes collapses or the stack falls over. Consequently, the relevance of studying creep behaviour of paper and board is that it can reduce stacking factors in design of corrugated board packages. This is a future goal in the development of a userfriendly computer-based tool for strength design of containers. Finally, this work shows how far it is possible to predict box performance using an orthotropic linear elastic material model, multi-ply eight node iso-parametric finite element and the Tsai-Wu failure criterion.
6
Presented papers
Paper 1:
T. Nordstrand, H. G. Allen and L. A. Carlsson, "Transverse Shear Stiffness of Structural Core Sandwich", Composite Structures, No. 27, pp. 317-329, 1994.
Paper 2:
T. Nordstrand and L. A. Carlsson, "Evaluation of Transverse Shear Stiffness of Structural Core Sandwich Plates", Composite Structures, Vol. 37, pp. 145153, 1997.
Paper 3:
T. Nordstrand, "On Buckling Loads for Edge-Loaded Orthotropic Plates including Transverse Shear". To be submitted to Composite Structures.
Paper 4:
T. Nordstrand, "Parametrical Study of the Post-buckling Strength of Structural Core Sandwich Panels", Composite Structures, Vol. 30, pp. 441451, 1995.
Paper 5:
T. Nordstrand, "Analysis and Testing of Corrugated Board Panels into the Post-buckling Regime". To be submitted to Composite Structures.
Paper 6:
P. Patel, T. Nordstrand and L. A. Carlsson, "Local buckling and collapse of corrugated board under biaxial stress", Composite Structures, Vol. 39, No. 12, pp. 93-110, 1997.
Paper 7:
T. Nordstrand, M. Blackenfeldt and M. Renman, "A Strength Prediction Method for Corrugated Board Containers", Report TVSM-3065, Div. of Structural Mechanics, Lund University, Sweden, 2003.
7
Acknowledgements
Firstly, I am indebted to the vision of the late Alf de Ruvo for initiating a far-reaching project at SCA called "Box Mechanics" in 1989. This project involved several staff researchers at SCA, and I would like to acknowledge the contributions from Dr. M. Blackenfeldt, Tekn.lic. M. Renman, Dr. P. Patel, Tekn. lic. Rickard Hägglund and M. Sc. Andreas Allansson. Secondly, the drive and support given by Dr. Leif Carlsson, Florida Atlantic University, has been invaluable for completing this task. I would also like to thank Dr. Per Johan Gustafsson, Lund University, for his guidance, comments and suggestions while completing this thesis. Finally, I would like to thank my family and friends for all their support. Sundsvall in December 2002, Tomas Nordstrand
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References
1.
R. Steadman, "Corrugated Board", Ch. 11, in Handbook of Physical Testing of Paper, (R.E. Mark et al. eds.), pp. 563-660, Marcel Dekker, New York, 2002.
2.
FEFCO, Avenue Louise 250, B – 1050 Brussels, Belgium, www.fefco.org.
3.
R. C. McKee and J. W. Gander, "Top-Load Compression", TAPPI, Vol. 40, No. 1, pp. 57-64, 1957.
4.
R. C. McKee, J. W. Cander, and J. R. Wachuta, "Compression Strength Formula for Corrugated Boxes", Paperboard Packaging, 48, pp. 149-159, 1963.
5.
M. W. Johnson and T. J. Urbanik, "Analysis of the Localized Buckling in Composite Plate Structures with Application to Determining the Strength of Corrugated Fiberboard", J. Comp. Tech. & Res., 11 (4), pp.121-127, 1989.
6.
B. S. Westerlind and L. A. Carlsson, "Compressive Response of Corrugated Board", TAPPI, 75 (7), pp.145-154, 1992.
7.
P. Patel, T. Nordstrand and L.A. Carlsson, "Instability and Failure of Corrugated Core Sandwich Cylinders Under Combined Stresses", in Multiaxial Fatigue and Deformation Testing Techniques, ASTM STP 1280(S. Kalluri and P.J. Bonacuse, Eds.), pp. 264-289, 1997.
8.
P. Patel, T. Nordstrand and L. A. Carlsson, "A Study on the Influence of Local Buckling on the Strength of Structural Core Sandwich Structures", Proceedings of EUROMECH 360, Ecole des Mines de Saint-Etienne, 1997.
9.
A. Allansson and B. Svärd, "Stability and Collapse of Corrugated Board", Master thesis, Report TVSM-5102, Division of Structural Mechanics, Lund University, Sweden, 2001.
10. E. K. Hahn, A.de Ruvo and L. A. Carlsson, " Compressive Strength of Edge-loaded Corrugated Panels", Exper. Mech., Vol. 32, pp. 252-258, 1992. 11. W. Vollmer, "Components of Compression Testing on Corrugated Board Packaging ", FEFCO. XIIIth Congress, 1974. 12. J. S. Buchanan, "The Effect of Crease Form on the Compressive Strength of Corrugated Cases", Packaging, pp. 37-43, March, 1963. 13. M. Renman, "A mechanical characterization of creased zones of corrugated board", Licentiate thesis, Report 94:05, Division of Engineering Logistics, Lund University, Sweden, 1994.
9
14. W. S. Peterson, "Unified Container Performance and Failure Theory I : Theoretical Development of Mathematical Model", TAPPI, Vol. 63(10), pp. 75-79, 1980. 15. J. C. Pommier and J. Poustis, "Bending Stiffness of Corrugated Board Prediction Using the Finite Element Method, Mechanics of Wood and Paper Materials", ASME AMD-Vol. 112, Edited by R.W. Perkins, pp. 67-70, 1990. 16. J. C. Pommier, J. Poustis,Bending, J. Fourcade and P. Morlier, "Determination of the critical load of a Corrugated Box Subjected to Vertical Compression by Finite Element Method", Proceedings of the 1991 International Paper Physics Conference, pp. 437-447, Kona, HI, 1991. 17. A. Rahman, "Finite element buckling analysis of corrugated fiberboard panels", Proceedings of the 1997 joint ASME/ASCE/SES summer meeting entitled mechanics of cellulosic materials, pp. 87-92; June 29-July 02, 1997. 18. P. Patel , "Biaxial Failure of Corrugated Board", Licentiate thesis, Division of Eng. Logistics, Lund University, Sweden 1996. 19. U. Nyman, P.J. Gustafsson, "Material and Structural Failure Criterion of Corrugated Board Facings", Accepted for publication in Composite Structures. 20. A. C. Gilchrist, J. C. Suhling and T. J. Urbanik., "Nonlinear finite element modeling of corrugated board", Mechanics of Cellulose Materials- ASME AMD-Vol 231/MDVol 85, pp. 101-106, 1999. 21. C. A. Bronkhorst and J. R. Riedemann, "The Creep Deformation Behaviour of Corrugated Containers in a Cyclic Moisture Environment", Proceedings of the Symposium on Moisture Induced Creep Behaviour of Paper and Board, pp. 249-273, Stockholm , Sweden, December 5-7, 1994. 22. T. Nordstrand and R. Hagglund, "Predicting Sagging of Corrugated Board Tray Bottom Using Isochrones", Proceedings of the 3 rd International Symposium on Moisture and Creep Effects on Paper, Board and Containers, pp. 215-220, Rotorua, New Zeeland, February 20-21, 1997.
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Part II Appended Papers
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Paper 1
_______
Transverse Shear Stiffness of Structural Core Sandwich by T. Nordstrand, Carlsson, L.A. and Allen, H.G. Composite Structures, Vol. 27, pp. 317-329, 1994.
Detta är en tom sida!
Paper 2
_______
Evaluation of Transverse Shear Stiffness of Structural Core Sandwich Plates by T. Nordstrand and Carlsson, L.A. Composite Structures, Vol. 37, pp. 145-153, 1997.
Detta är en tom sida!
Paper 3
_______
On Buckling Loads for Edge-Loaded Orthotropic Plates including Transverse Shear by T. Nordstrand To be submitted to Composite Structures.
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On Buckling Loads for Edge-Loaded Orthotropic Plates including Transverse Shear Tomas Nordstrand SCA Research, Box 716, 851 21 Sundsvall, Sweden
ABSTRACT
Corrugated board usually exhibits low transverse shear stiffness, especially across the corrugations. In the present study the transverse shear is included in an analysis to predict the critical buckling load of an edge-loaded orthotropic linear elastic sandwich plate with all edges simply supported. In the analysis, effective (homogenised) properties of the corrugated core are used. Classical elastic buckling theory of orthotropic sandwich plates predicts that such plates have a finite buckling coefficient when the aspect ratio, i.e. the ratio between the height and width of the plate, becomes small. However, inclusion in the governing equilibrium equations of the additional moments, produced by the membrane stresses in the plate at large transverse shear deformations, gives a buckling coefficient which approaches infinity when the aspect ratio goes to zero. This improvement was first included in the buckling theory of helical springs by Harinx (1942) and later applied to orthotropic plates by Burt and Chang (1972). Some inconsistencies in the latter analysis have been considered. The critical buckling load calculated with corrected analysis is compared with a predicted load obtained using finite element analysis of a corrugated board panel, and also with the critical buckling load obtained from panel compression tests.
1
INTRODUCTION
Corrugated board usually exhibits low transverse shear stiffness, especially across corrugations[1, 2]. This will reduce the critical buckling load according to classical theory of orthotropic sandwich panels [3, 4]. In this small-deflection theory it is customary to assume that the membrane forces are unchanged during plate deflection and equal to their initial values. However, due to the large transverse shear strains, the change in direction of the membrane forces over a small plate element can not be disregarded. This gives additional moments that are introduced in the governing moment equilibrium equations of the panel. Such additional moments were first included in the buckling theory of helical springs by Harinx [5]. Later this was applied to shear deformable plates by Bert and Chang [6] although their work contains some inconsistencies that are corrected herein. Furthermore, in the corrected analysis the expression for the buckling coefficient is shown to reduce to the classical formulation of an orthotropic plate without shear deformation when the transverse shear stiffnesses become large. It is also shown that the buckling coefficient goes to infinity when the height-width ratio of the plate is decreased towards zero. In the following analysis the corrugated board panel is regarded as a laminated shear deformable orthotropic linear elastic plate[7]. Thus, effective (homogenised) properties of the corrugated core are used [8, 9]. The papers in the facings are also regarded as orthotropic linear elastic materials [10,11]. The analysis was used to confirm predicted critical buckling load from a finite element analysis of a corrugated board panel modelled with eight-node multi-layered isoparametric shell elements [12-14]. Predicted critical buckling load is also compared to buckling loads obtained from compression tests of corrugated board panels [15].
ANALYSIS
Figure 1 shows an element of a corrugated board panel of thickness h. Core height is hc and wavelength of corrugations is λ M. Facing thickness is t f and thickness of core sheet is t c. The principal axes of elastic symmetry of the face sheets and the core are aligned with the Cartesian coordinate system xyz. The 2-axes of the corrugated medium is parallel with the y-axes.
z tf
h hc t c
3
y,CD 2
λ M
1 b
x,MD
Figure 1. Schematic diagram diagram of corrugated corrugated board. It is assumed that the facings and core sheet are thin compared to the total thickness of the panel and that the transverse shear strains are uniform in the core layer.
2
Furthermore, the deflections and slopes are assumed to be small compared to the thickness of the plate. Transverse shear deformation of the plate is accommodated by assuming that cross-sections remains straight but not necessarily normal to the midplane of the plate during d uring bending [6]. The membrane forces N x , N y , N xy , transverse shear forces Q xz , Q yz , bending and twisting moments M x , M y , M xy are acting on respective four sides of a plate element, see Figure 2. The transverse shear strains, see Fig. 3, are determined by [4]
γ xz = A55Q xz
(1)
γ yz = A44 Q yz where A44 and A55 are the transverse shear stiffnesses [1, 2]. x dx z y
dy Mxdy
Mxydy Mydx
Nxdy
Mxydx
h
Nxydx
Nxydy
Qxdy Nydx
Figure 2.
Qydx
Forces and moments acting on a plate element hdydx.
The plate displacement w is then related to the applied moments as follows [4] M x = − D11 M y = − D12
∂ β x ∂ x ∂ β x ∂ x
1
where
∂ β y
− D12
∂ y
− D 22
∂ β
∂ β y ∂ y
∂ β y
M xy = − D66 x + 2 y x ∂ ∂ ∂w ∂w
β x =
∂ x
− γ xz ,
β y =
∂ y
− γ yz
(2)
(3)
and D11, D22, D12 and D66 are the bending and twisting stiffnesses defined according to ref. [7-11]. 3
Figure 3 shows a cross-sectional view of the deformed plate element in the x-z-plane. Considering that the plate is loaded in compression, the normal forces N x, Ny and the shear force N xy are much larger than the transverse shear forces Q yz and Qxz and have to be accounted for in the lateral equilibrium of the differential plate element. Subsequently, after algebraic manipulation the equation of equilibrium in the zdirection is obtained as,
∂β y ∂β ∂β y ∂ Q xz ∂ Q yz ∂β = 0 + Nx x + Ny + + 2N xy x + ∂ x ∂ y ∂ x ∂ y ∂ ∂ y x
y
dx
.
x
∂w ∂x β w
(4)
x
Q
Q + xz
xz
∂Q
γ N xz x M
xz dx x ∂
M + x
x
∂M
x
∂ x
dx
∂ γ xz ∂x ∂β
z
Figure 3.
N + x
∂ Nx ∂x
2 ∂ w ∂x dx + ∂x ∂x 2
∂w dx
x+
∂ 2γ xz 2
∂x
x dx
2
x+
∂β ∂x
2
x dx
Cross-section of a differential plate element.
The transverse shear strains at the left and right cross-sections in Figure 3 reduce the slope slightly more of the right cross-section than the left cross-sections. This will rotate the normal forces so that their action will not be through the centre of the differential plate element. Consequently, the normal forces will generate additional moments. These moments are taken into account in the present theory. This is the basic difference between the present theory and the classical sandwich theory [4]. The derivation of moment equilibrium around an axis through the centre of the differential element and parallel with the y-axes in Fig. 3 is then as follows.
∂ M yx ∂ M x M + dx dy M dy + M dx M + dy dx x x yx yx ∂ ∂ x y ∂ Q xz ∂ N x ∂γ - Q xz + dx dydx + N x + dx dydx γ xz + xz dx = 0 ∂ x ∂ x ∂ x
(5)
If bot both h sid sides es in eq. eq. (5) (5) are are div diviided ded by by dx dxdy and and let lettting ing dx → 0 and and dy → 0 , eq. (5) reduces to
∂M x ∂M xy - Q xz + N x γ xz xz = 0 ∂x ∂y
(6a)
4
Similarly, moment equilibrium around an axis through the centre of the differential element parallel with the x-axis yields
∂M y ∂M yx - Q yz + N y γ yz = 0 ∂y ∂x
(6b)
Substitution of eqs. (1)-(3) in equilibrium eqs. (4) and (6) forms a system of three simultaneous differential equations in terms of the out-of-plane displacement w and the transverse shear strains γ xz and γ yz . It is assumed that the plate is simply supported along its edges, i.e. the edges of the panel are prevented from moving out-of-plane and are not rotationally restrained. The edges are also free to move in-plane and transverse shear strains are prevented by edge stiffeners. The edges of the panel, parallel to the x-axis, are compressed uniformly by a load of intensity, p y, per unit length, see Fig. 4.
y a
z
py
b x
Figure 4. Schematic diagram of a simply supported panel in edgewise compression. Thus, boundary conditions at y=0 and y=b are w=0
My = 0
γ xz = 0
Mxy = 0
and boundary conditions at x = 0 and x = a are w=0
Mx = 0
γ yz = 0
Mxy = 0
According to Navier´s procedure [7], a solution of the three simultaneous differential equations that satisfy the boundary conditions above can be obtained by assuming that the out-of-plane displacement, w, and transverse shear strains γ xz, γ yz can be represented by double trigonometric series. However, in the present analysis only oneterm solutions are used for w, γ xz and γ yz , respectively. w = W sin(Ax ) sin(By) xz
(7a)
= Γ xz cos( Ax) sin( By )
(7b)
5
γ yz = Γ yz sin( Ax) cos( By )
(7c)
where W , Γ xz and Γ yz are corresponding amplitudes. A and B are A=
mπ a
, B=
nπ b
(8)
Integers m and n are number of buckles, i.e. m and n half sine waves, in the x and y directions. In the subsequent analysis it is convenient to define a number of parameters of the homogenised sandwich plate [4]: 2
na λ = , ζ = mb
D11 , ψ = D 22
D 66 D11D 22
, η=
D12 + 2D 66 D11D 22
, s xz =
A 55 k py
, s yz =
A 44 k py
where
a k = nπ
2
p y,crit D11D 22
, Pcrit ,theor = a p y ,crit
(9)
py,crit , is the critical load intensity (load/unit length) to cause panel buckling. The total critical buckling load is P crit,theor. Substitution of the trigonometric expressions eq. (7) in the differential equations (4), (6a) and (6b) leads to the following system of equations shown in matrix form 1 λ - (k + s yz ) - s xz k A B 1 λ 1 + λη - 1 + ψ + k + s yz - (η - ψ ) ζ A λζ B 1 λ ψ - (η - ψ ) - ζ + + s xz λζ + η λ A B
W 0 Γ xz = 0 Γ 0 yz
(10)
Correct expressions of the elements in the stiffness matrix of eq. (10) are given instead of those in the stiffness formulation [6]. Solution of eqs. (10) different than the trivial one, W = Γ xz = Γ yz = 0 are possible when the determinant of the matrix vanishes. This criterion leads to a second order equation of k
Pk 2 + Qk + R = 0
(11)
where
6
P = ζ +
ψ + s xz λ
(12a)
Q = (s yz + ψ - λζ - 2η )P +
(λζ + η )2 - (η - ψ )2
(12b)
λ
1 (λζ + η )2 ψ − PΘ - s yz P λζ + 2η + R = ζ + Θ + s yz λ λ λζ
(12c)
and
1 2 - η Θ = 1 + ψ λζ + 2η + λζ
(12d)
The non-trivial solution of eq. (8) can thus be found when Q k = − ± 2P
Q2 4P
2
−
R P
(13)
where only positive values of k are valid since the buckling load must be compressive. The critical buckling load, P crit,theor, is given by eq.(9), where n=1 and k is the smallest positive value given by eq. (11). Using eq. (12), the two ratios in eq. (13) are: 2 2 s ψ - λζ - 2η (λζ + η ) - (η - ψ ) Q yz = + 1 + 2P 2 s yz ψ ζ + λ s xz s yz λ 1 + s xz
(14a)
1 2 ψ λζ η η 1 + + 2 + λζ R 1 (λζ + η )2 + + = -s yz λζ + 2η + λζ P ψ ψ ζ ζ + + λ λ s xz λ 1 + s yz 1 + s xz s xz
(14b)
Attention is now turned to analysis of the limit case of infinite large transverse shear stiffnesses in order to show that the buckling coefficient k, determined by eq. (13), for that limit is reduced to the buckling coefficient for orthotropic plates without shear deformation [7].
7
If the transverse shear stiffnesses A 44 and A 55 , corresponding to s yz and s xz , approach infinity then it is evident from eqs. (14) that s yz Q → 2P 2
(15a)
and
R 1 → −s yz λζ + 2η + P λζ
(15b)
Substitution of eqs. (15) into eq. (13) gives
1 λζ η 4 +2 + s yz λζ − k = − 1 1 + 2 s yz
(16)
The square root in eq. (16) can be expanded according to the binomial series 1+ χ = 1 +
1 1 χ − χ 2 +..... 2 4
(17)
where
4 λζ + 2η +
χ =
1 λζ
(18)
s yz
If eq. (17) and eq. (18) is substituted into eq. (16), the expression on the right hand side is reduced to the buckling coefficient for an orthotropic plate [4] k = λζ + 2 η +
1
(19)
λζ
when s yz goes to infinity.
8
10 k , t n e i c i f f e o c g n i l k c u B
9
With Shear
8
Without Shear
7
Sandwich theory
6 5 4 3 2 0
0,5
1
1,5
2
Plate height/width, b/a Figure 5.
Buckling coefficient k, according to present theory, eq.(13)-(14), the theory for orthotropic plate without shear [4] and the classical sandwich theory [4].
In Figure 5 the buckling coefficient for a plate with and without transverse shear is plotted versus the plate width/height ratio. The material data used is typical for a common corrugated board grade and defined in Table 2. Notice that the buckling coefficient of the plate including transverse shear has no limit when the plate height/width becomes small as classical sandwich buckling theory predicts [3,4]. COMPARISON WITH FINITE ELEMENT ANALYSIS OF A CORRUGATED BOARD PANEL
In a finite element analysis of a simply supported corrugated board panel, with side lengths a = b = 400 mm, following eigenvalue analysis was made to obtain the critical buckling load [12]. A multi-ply eight node isoparametric shell element where first order transverse shear deformation is accounted for is used in the analysis. A quarter of the panel was modelled, due to symmetry, in a 6 x6 element mesh. The side length ratio between the corner element and mid element was 1:5. The finite element eigenvalue analysis is
([K ] + χ [S]ref ){ψ } = {0}
(20)
where [K] is the global stiffness matrix of the finite element model, S
ref
is the
"stress stiffness matrix", χ is the factor used to multiply the loads which generate the stresses and {ψ } is the generalised displacement vector of the nodes [13]. The load
9
{P} is also scaled by χ and it alters the intensity of the membrane stresses but not the distribution of the stresses such that
{P} = {P}ref ⇔ [S] = χ [S]ref
(21)
As χ is increased, the overall stiffness of the plate, ([Κ] + [S]), is reduced until a critical load {P}cr corresponding to the eigenvalue χcr is reached and the plate becomes unstable, i.e. det ([Κ] + [S]) goes to zero. The corrugated board analysed has 0.23 mm thick liners and a corrugated medium with wall thickness 0.25 mm and wavelength 7.26 mm. The height of the core layer is h c = 3.65 mm, see Fig. 1. Using the material data in Table 1, the buckling load of the corrugated board panel was calculated to P cr,fem = 849 N. This value is in excellent agreement with the value obtained by the closed form solution P cr,theor = 846 N, see eq. (9) and eq. (13). Sandwich theory gives a critical buckling P cr,sand = 815 N and an orthotropic plate without shear P cr,ortho = 898 N. Table 1. Effective material properties of the layers in the panel. Layer 1 2 3 Layer 1 2 3 Layer
Ex (GPa) 8.25 0.005 8.18 Gxy (GPa) 1.89 0.005 1.95
1 2 3
νxy
Ey (GPa) 2.9 0.231 3.12 Gxz (GPa) 0.007 0.0035 0.007 νxz +
Ez (GPa) 2.9 3.0 3.12 Gyz (GPa) 0.070 0.035 0.070 νyz +
0.43 0.05 0.43
0.01 0.01 0.01
0.01 0.01 0.01
+ The Poisson's ratios are assumed small because of the plane stress condition in the board.
COMPARISON WITH EXPERIMENTS
Panels size 400 x400 mm were cut from corrugated board and tested under compression in a rig that furnishes simply supported boundary conditions [15]. Panels were oriented with the cross direction (CD) in the direction of loading, see Figure 6.
10
Figure 6.
Rig and corrugated board panel tested under compression.
Material data for the board is given in Table 2. The critical buckling load as estimated from the test results by means of a non-linear regression analysis method [15] was 814 N. This value is consistent with the analytically predicted critical buckling load of 870 N using the present buckling analysis, i.e. an analysis of a plate including transverse shear deformation. Table 2. Corrugated board data. Transverse shear stiffness is measured. Basis weight, g/m 2 Thickness, mm Corrugation wavelength, mm Bending stiffness, Nm
Transverse shear stiffness, kN/m
h
λ M D 11 D22 D12 D66 A 44 A55
556 4.02 7.26 14.6 5.43 2.71 3.34 39.2 5.6
In comparison, a plate without shear deformation is predicted to have a critical buckling load of 924 N, which is exactly the same result as obtained from the classical theory of orthotropic plates [7].
CONCLUSIONS
An explicit equation for the buckling load of a simply supported orthotropic linear elastic plate in edgewise compression has been derived taking into account first order transverse shear deformation. There is major difference between present theory and classical sandwich theory in the additional moments that are introduced in the governing moment equilibrium equations of the panel, due to change in directions of the membrane forces over a small plate element that has large transverse shear strains.
11
When the transverse shear stiffness goes to infinity the critical buckling load, predicted by the present theory, is shown to be reduced to the critical buckling load of an orthotropic plate without transverse shear deformation. Furthermore, the buckling coefficient does not have a limit in the present theory when the plate height/width becomes small, as classical sandwich buckling theory predicts. The present theory is approximate due to one-term approximations of the deflection w(x,y) and the transverse shear strains γ xz(x,y) and γ yz(x,y). Verification by finite element analysis suggests that the present explicit equation for the buckling load is accurate, the deviation is typically less than 0.5%. However, the discrepancy is larger between present theoretical buckling load and the experimental buckling load of corrugated board panels. This may partly be due to the difficulties involved in evaluation of the buckling load from the experimental results [15] partly due to the non-linear material behaviour of paper.
REFERENCES
1.
T. M. Nordstrand, H. G. Allen and L. A. Carlsson, "Transverse Shear Stiffness of Structural Core Sandwich", Composite Structures, Vol. 27, pp. 317-329, 1994.
2.
T. Nordstrand and L.A. Carlsson, “Evaluation of Transverse Shear Stiffness of Structural Core Sandwich Plates”, Composite Structures, Vol. 37, pp. 145-153, 1997.
3.
E. Reissner, “The Effect of Transverse Shear Deformation on Bending of Elastic Plates”, Journal of Applied Mechanics, 12, A69-A77, 1945.
4.
F.J. Plantema, “Sandwich Construction”, John Wiley & Sons, Inc., New York, N.Y., 1966.
5.
J. A. Harinx, “On Buckling and the Lateral Rigidity of Helical Compression Springs”, Proceedings of the Konjlike Nederland Akademie Wettenschappen, Vol. 45, pp. 533-539, 650-654, Amsterdam, Holland,1942.
6.
C.W. Bert and S. Chang, “Shear-Flexible Orthotropic Plates Loaded In-Plane”, Journal of Engineering Mechanics Division, Proceedings of the American Society of Civil Engineers, Vol. 98, No. EM6, pp. 1499-1509, December, 1972.
7.
Jones, R. M., Mechanics of Composite Materials, Hemisphere Publishing Corp. New York, 1975.
8.
L. A. Carlsson, T. Nordstrand and B. Westerlind, “On the Elastic Stiffnesses of Corrugated Core Sandwich”, Journal of Sandwich Sturctures and Materials, Vol. 3, pp. 253-267, 2001.
9.
C. Libove and R. E. Hubka, "Elastic Constants for Corrugated Core Sandwich Plates", NACA, TN 2289, 1951.
10. G. A. Baum, D. C. Brennan and C. C. Habeger, "Orthotropic Elastic Constants of Paper", Tappi, Vol. 64, pp. 97-101, 1981. 11. R. Paetow and L. Göttsching, “Poisson’s Ratio of Paper”, Das Papier, Vol. 6, pp. 229-237, 1990.
12
12. T. Nordstrand, “Parametrical Study of the Post-buckling Strength of Structural Core Sandwich Panels”, Composite Structures, Vol. 30, pp. 441-451, 1995. 13. ANSYS User’s Manual, Swanson Analysis System Inc., Vol. 1-2, 1989. 14. S. Ahmad, B. M. Irons and O. C. Zienkiewicz, ”Analysis of Thick and Thin Shell Structures by Curved Finite Elements”, International Journal of Numerical Methods in Engineering, No. 2, pp 419-451, 1970. 15. T. Nordstrand, “Analysis and Testing of Corrugated Board Panels into the Postbuckling Regime”, To be submitted to Composite Structures.
13
Detta är en tom sida!
Paper 4
_______
Parametrical Study of the Post-buckling Strength of Structural Core Sandwich Panels by T. Nordstrand Composite Structures, Vol. 30, pp. 441-451, 1995.
Detta är en tom sida!
Paper 5
_______
Analysis and Testing of Corrugated Board Panels into the Post-buckling Regime by T. Nordstrand To be submitted to Composite Structures.
Detta är en tom sida!
Analysis and Testing of Corrugated Board Panels into the Post-buckling Regime Tomas Nordstrand SCA Research, Box 716, 851 21 Sundsvall, Sweden
ABSTRACT
Testing of the load bearing capacity of corrugated board boxes is often associated with uncertainties, e.g. the creases along the edges of the side panels introduce eccentricities. An alternative to the testing of boxes is therefore attractive. One suggestion is testing of panels. However, panels are sensitive to the boundary conditions. A panel compression test (PCT-) rig, similar to a test frame for metal plates designed by A. C. Walker, was therefore built to achieve accurately defined load and boundary conditions. The PCT-rig furnishes simply supported boundary conditions, i.e. the edges of the panel are prevented from moving out-of-plane without any rotational restraint. The edges are also free to move in-plane. In order to describe the buckling behaviour, a non-linear buckling analysis of orthotropic plates, derived by Banks and Harvey, was modified to include initial imperfections. The critical buckling load of the panels was evaluated by fitting the analytical expression by non-linear regression to experimentally measured load-displacement curves. The results show a difference in the order of 15-20 % between experimentally estimated critical buckling load and the analytically predicted critical buckling load for orthotropic plates. This is mainly attributed to transverse shear deformations. A corresponding difference was observed between analytically predicted and experimentally measured load-displacement curves at large out-of-plane deformation, i.e. twice or three times the board thickness. This is probably caused by the non-linear response of paper at high stresses and local buckling of the panel facings, i.e. the liners. A predicted failure load of the corrugated board panel was determined when stresses in the facings reached the Tsai-Wu failure criterion. The predicted failure load and measured average experimental failure load were close, indicating that collapse of the panel is triggered by material failure of one of the liners.
NOTATIONS
A, A0
Plate size in x and y directions Amplitudes of the total and initial plate deflection functions
Ai
Relative amplitude of the shape function
a, b
Dij
Bending stiffness of the plate
Eij
Modulus of elasticity of liner
Eij,c
Modulus of elasticity of medium
Gij
Shear modulus of liner
F
Airy’s stress function
h
Plate thickness
hc
Core thickness
t f
Liner thickness
t c α
Medium thickness Take-up factor
M x, M y, M xy
Bending and twisting moments per unit distance in middle surface of the plate
N x, N y, N xy
Membrane forces per unit distance in middle surface of the plate
P
Load
Pcrit
Critical buckling load
u, v, w
Displacements in x, y and z directions
V
Strain energy
x, y, z
Cartesian co-ordinates
γ
Unit shear strain
ε x , ε y
Unit normal strains in x and y directions of the facings
ν ij
Poisson’s ratio
σ x , σ y
Unit normal stresses in x and y directions
τ xy
Unit shear stress on plane perpendicular to the x-axis and parallel to the y-axis
INTRODUCTION
Corrugated board is one of our most common transport packaging materials. Large retailers and distributors are under increasing pressure to cut the cost of corrugated packaging. With the increasing scale of business it has become unacceptable to over design boxes. Consequently, it is necessary to predict box strength in order to obtain boxes at the lowest possible cost. However, analysis of top-to-bottom compression loading of boxes is often associated with uncertainties, e.g. the creases between flaps and side panels introduce eccentricities along the loaded edges [1]. Since the buckling behaviour is of primary interest, it was decided to test corrugated board panels with clean cut edges in a specially designed panel compression rig, similar to a test frame for metal plates [2]. The panel compression rig furnishes simply supported boundary conditions, i.e. the edges of the panel are prevented from moving out-of-plane without any rotational restraint. The edges are also free to move in-plane. It was decided to measure the outof-plane displacement at the centre of the panel versus the compressive load. Out-ofplane measurement of the panel deformation is easier than in-plane measurement. It also simplifies (de-)mounting of the panel in the rig. One objective of the tests is to obtain the critical buckling load. Since post-buckling of a panel is stable, an analytical expression was needed that relates the compressive load to the deformation of the panel. Banks and Harvey [3] originally derived a postbuckling analysis, which has been modified in the presented model to include initial imperfections. Panels are assumed to have orthotropic elastic constants as described by Jones [4]. The critical buckling load of the panels was evaluated by fitting an analytical expression for the load-deformation curve to the experimentally measured curves. The fitting was made by non-linear regression analysis and comprised the determination of three parameters in the analytical expression, one being the buckling load, another the post-buckling coefficient and the third the amplitude of the initial imperfection of the panel. Results show a discrepancy of 15-20 % between experimentally estimated critical buckling load and the theoretically predicted buckling load for orthotropic plates. This difference is mainly attributed to transverse shear. A corresponding load difference was observed between analytically predicted and experimentally measured postbuckling curves at large deflections, i.e. twice or three times the board thickness. This is probably caused by the non-linear material response of paper at high stresses and by local buckling of the panel facings, i.e. the liners. A failure load of the corrugated board panel was predicted by determining when stresses in the facings reached the Tsai-Wu failure criterion [5]. The predicted failure load and measured average experimental failure load were close, indicating that collapse of the panel is triggered by material failure of one of the liners. Thus, the strength of the material is efficiently utilised.
1
THEORETICAL MODEL Basic assumptions and mechanics principles of a corrugated board panel
A simply supported corrugated board panel, loaded in compression, buckles in a stable manner and carries load beyond the critical buckling load until compressive failure occurs. Since the paper sheets used in the panel are thin compared to the overall thickness of the panel, the variations of stresses in the thickness direction of each sheet are ignored. Paper Roll 2
Panel
1
z
tf
h c tc
y,CD 2 1
λ
x,MD
Figure 1. Geometry and orientation of a symmetrical corrugated board panel. The paper is assumed to have orthotropic elastic properties with the elastic planes of symmetry of the facings coinciding with the Cartesian coordinate system xyz of the panel, see Figure 1. Total thickness of the panel is h, core height is c and facings and core sheet are assumed to have thickness tf and tc, respectively. The membrane forces N x, Ny and Nxy are shown in Figure 2. These forces are oriented according to the orientation of the panel in the loaded state. The displacements are assumed to be small in the sense that sin (∂w / ∂x i ) = ∂w / ∂x i , i = 1, 2 and the projected membrane forces in the x-y plane are in equilibrium. z N Nxy dx
x
Nyx dy
N
Nx + N xy +
y
Ny +
∂Ny ∂ y
N yx +
dy
∂N yx ∂ y
∂N xy ∂ x
dx
dy
Figure 2. Membrane forces in the corrugated board panel.
2
∂Nx dx ∂ x
It is assumed that the membrane strains are constant through the thickness of the panel and that membrane forces carried by the corrugated core in x-direction can be disregarded. The strains in panel facings due to membrane forces are
h ε x 2t f E 11 h ε ν = − h 12 y 2t f E 11 γ xy 0 where
h
1
= 1+
α t c E 22,c 2t f
−ν 21
N x h N y 0 N h h xy 2t f G12 h
hh
0
2t f E 22 hh
2t f E 22 0
(1)
and ν ij , E ij i,j =1,2 and G12 are the Poisson´s ratio, elastic and
E 22
shear modulus of the facings, respectively. E 22,c is the elastic modulus of the core sheet in the cross direction CD and α is the take-up factor, i.e. the ratio between the length of the corrugated core sheet and the length of the board.
z
My Mxy
dx
Myx
x
dy
Mx
Mx + M xy +
y
My +
Figure 3.
∂M y ∂ y
M yx +
dy
∂M yx ∂ y
∂M x dx ∂ x
∂M xy ∂ x
dx
dy
Bending of the corrugated board panel.
The bending and twisting curvatures and strains in the facings z = ±
κ x
ε x (z) − ε x (0) ∂2w = =− 2 , z ∂x
κ y
=
ε y (z ) − ε y (0 ) z
=−
∂2w , ∂y 2
κ xy
=
xy
h − t f
2
(z ) − γ xy (0 ) z
= −2
∂2w ∂x∂y
(2)
are connected to the bending and twisting moments acting on the panel in Figure 3 by
M x D11 M y = D12 M xy 0
D12 D 22 0
0 κ x
0 κ y D 66 κ xy
where D ij are elements of the bending stiffness matrix [4,5]. 3
(3)
Buckling of a corrugated board panel
When a panel with a small initial curvature is subjected to a compressive load it bends, and the deflection may become large in comparison with the thickness of the board, see Figure 4. Hence the geometrical non-linearity due to membrane stretching has to be included in the analysis [6]. It is assumed that the membrane strains are constant through the thickness of the panel, and can be expressed in terms of the displacements at z=0 as follows
∂u 1 ∂w ε x = + ∂ x 2 ∂ x
2
2
∂v 1 ∂w ε y = + ∂ y 2 ∂ y ∂u ∂v ∂w ∂w γ xy = + + ∂ y ∂ x ∂ x ∂ y
(4)
By adding the second derivative of εx with respect of y and second derivative of
εy
with respect of x and subtracting the second derivative of γ xy with respect of x and y, a compatibility relation is obtained between the membrane strains and the out-ofplane displacement [6]. 2
2 2 ∂ 2 ε x ∂ ε y ∂ γ xy ∂ 2 w ∂2w ∂2w + 2 − = − ∂ x 2 ∂ y 2 ∂ y 2 ∂x ∂x∂y ∂ x∂y
(5)
If the expression in the stiffness matrix is substituted with the effective elastic stiffnesses of the panel as follows * = E 11
2t f h
* = E 11 , E 22
2t f hh
* = E 22 , G12
2t f h
* = hν 12 G12 , ν 12
(6)
and the strains are subsequently substituted in eq. (2) we obtain 2
2 2ν 12* ∂ 2 N xy 1 ∂ N y 1 1 1 ∂ 2 N x ∂ 2 w ∂ 2 w ∂ 2 w − 2 − + + * = 2 hE *22 ∂ x 2 h G12* E 11* ∂ x∂ y hE 11 ∂ y 2 x y ∂ ∂ ∂ x ∂ y
(7)
The solution of eq. (7) can be greatly simplified by introducing Airy's stress function, F [6]. With this stress function the membrane forces in eq. (7) can be expressed as
∂ 2 F ∂ 2 F ∂ 2 F , N y = h 2 , N xy = −h N x = h ∂ x∂ y ∂ y 2 ∂ x where F=F(x,y). With these expressions for the forces, eq. (7) becomes
4
(8)
2
4 2 2 2 2ν 12* ∂ 4 F 1 ∂ 4 F 1 1 F w w w ∂ ∂ ∂ ∂ + − + = − E *22 ∂ x 4 G12* E 11* ∂ x 2 ∂ y 2 E 11* ∂ y 4 ∂ x∂ y ∂ x 2 ∂ y 2
(9)
If w=w(x,y) is the total out-of-plane displacement and w0 = w0 (x,y) is an initial imperfection of the plate, eq. (9) can be written as 4 2ν 12* ∂ 4 F F 1 ∂ 4 F 1 1 ∂ + − + = E *22 ∂ x 4 G12* E 11* ∂ x 2 ∂ y 2 E 11* ∂ y 4 2 2 ∂ 2 w ∂ 2 w ∂ 2 w ∂ 2 w0 ∂ 2 w0 ∂ 2 w0 − 2 − 2 − 2 2 x y x y x y x y ∂ ∂ ∂ ∂ ∂ ∂ ∂ ∂
(10)
provided that the total displacements and the initial imperfections have the same shape, differing only in magnitude. This equation links the membrane stresses with out-of-plane displacements of orthotropic plates.
y
P
X
b
-∆ /2
z
∆ /2
P a
Figure 4. Simply supported corrugated board panel with compression of top and bottom edges. Boundary conditions
It is assumed that the panel is simply supported, see Figure 4. This means that edges are rotationally unrestrained and no out-of-plane displacement is present. Neither are in-plane shear stresses allowed. Thus at the unloaded edges x=0, a :
5
w=0
∂ 2 w ∂ 2 w M x = − D11 2 + ν 21 2 = 0 ∂ y ∂ x τ xy = 0 σx
(11)
=0
The compressive displacements are constant along the loaded edges. Thus, at y = ±
b
2
,
the conditions are w=0
∂ 2 w ∂ 2 w =0 M y = − D22 2 + ν 12 2 ∂ ∂ y x τ xy
=0
v=m
(12)
∆ 2
Solution strategy
In order to obtain a relationship between the compression v and out-of-plane displacement w the principle of minimum potential energy is used [3]. The total potential strain energy V in the buckled plate consists of two parts - the potential energy of bending and twisting, V B , and the membrane strain energy, V M . V = V B + V M b a 2
2 ∂ 2 ∂ 2 ∂ 2 1 D11 2 (w − w0 ) + 2 D11ν 21 2 (w − w0 ) 2 (w − w0 ) + V B = 2 0 b ∂ x ∂ x ∂ y
∫∫ −
2
2
∂ 2 ∂ 2 + D22 2 ( w − w0 ) + 4 D66 ( w − w0 ) ∂ y ∂ x∂y
2
dydx
(13)
b a 2
1 ∂ 2 F 2 2ν ∂ 2 F ∂ 2 F 1 ∂ 2 F 2 1 ∂ 2 F 2 2 − 12 2 + + dydx V M = ∫ ∫ 2 0 b E 11 ∂ y E 11 ∂ y ∂ x 2 G12 ∂ x∂ y E 22 ∂ x 2 − h
2
Before minimising the total energy V , the out-of-plane displacement w and the initial imperfection w 0 are prescribed as follows [3]
π y b
w = AX(x)cos
(14)
π y b
w 0 = A 0 X(x)cos
(15)
6
where A and A0 are the amplitude of the out-of-plane displacement and initial imperfection, respectively. The panel is accordingly assumed to have a sinusoidal shape in y-direction and X ( x) is a polynomial
x x 2 x 4 X = 3.2 − 2 + a a a
(16)
that describes the deflected shape of the panel in x-direction, see Figure 5. 1.5 1 0.5 0 0
Figure 5.
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
Shape function used to model the buckled plate.
Integration of the membrane deformation ∂v / ∂y in eq.(4) is equal to the uniform compression of the panel given by eq.(12) b
2 2 1 ∂w 1 ∂w 0 + dy v = ∫ ε y − ∂ ∂ 2 y 2 y b − 2
(17)
2
Combining eq.(1), (6) and (8) gives
∂ 2 F 1 ∂ 2 F ε y = * 2 − ν 21 2 E 22 ∂ x ∂ y
(18)
Assuming a solution to eq.(10) [3] for a panel with uniform compression of its edges as shown in Figure 4
2π y b
F = F 1 ( x ) + F 2 ( x ) cos
(19)
Substitution of eq. (14), (15), (18) and (19) into eq.(17) and integrating gives 2
* ∂ 2 F 1 E 22 π * v 2 2 2 = E 22 + ( A − A0 ) X 2 ∂ x 4 b b
(20)
Substitution of eq.(14), (15), (19) and (20) into eqs. (13) and minimising with respect to A yields the relationship between A and v C A0 C 1 − ( A 2 − A02 ) 3 C 1 A C 2
v = −1 −
(21)
where constants C 1 , C2 and C3 are presented in the Appendix with a detailed solution. 7
Substitution of eq.(21) into (20) the relationship between the applied load P and outof-plane displacement A is given by integrating the stress σy over the loaded edge y=b/2 a
a
a
∂ 2 F 1 ∂ 2 F P = − h ∫ σ y dx = −h ∫ 2 dx = −h ∫ dx 2 ∂ ∂ x x 0 0 0
(22)
Further integration of eq.(22) gives
P = Pcrit 1 −
A0
+ Ψ ( A 2 − A0 A
2
)
(23)
where the critical buckling load * Pcrit = E 22
ah C 2
(24)
b C 1
and the post-buckling parameter is (25)
E * ah C 3 Ψ= − 2C 1 b C 1 22
Failure criterion
From the solution above, the total stresses in the inner facing z = −(h − t f ) / 2 of the panel in Figure 4
∂ 2 ∂2 ( ) ( ) + − + − ν 21 σ x = w w w w 0 0 ∂ y 2 2t f ∂ y 2 (h − t f )t f ∂ x 2 h ∂ 2 F
σ
τ
∂ 2 ∂2 ( ) ( ) = + − + − ν w w w w 0 12 0 ∂ x 2 2t f ∂ x 2 (h − t f )t f ∂ y 2 hh ∂ 2 F
y
D11
D22
(26)
2 D66 ∂ 2 (w − w0 ) =− + 2t f ∂ x∂ y ( h − t f )t f ∂ x∂ y h ∂ 2 F
xy
can be obtained for a specified value of the out-of-plane deformation A. Subsequently, the stresses in eq. (26) are inserted in the Tsai-Wu failure criterion assuming plane stress [5]
Γ 1σ x + Γ 2σ y + Γ 11σ x2 + Γ 22σ y2 + Γ 66τ xy2 + 2Γ 12σ xσ y = 1
8
(27)
where Γ 11 = −
Γ 2 =
1 1 1 1 , Γ 22 = − , Γ 12 = − 0.36 F11F22 , Γ 1 = + , σ x,tσ x,c σ y,tσ y,c σ x,t σ x,c
1 1 1 + , Γ 66 = σ y,t σ y,c σ x,cσ y,c
The subscript j=t or j=c of strength σ i,j ,i=x,y, denote the strength in tension and compression, respectively. The expressions for Γ 12 and Γ 66 are approximations for paper materials [7]. A geometrical interpretation of the failure criterion, eq. (27), is depicted in Figure 6. Failure occurs when the total stress vector [ σx, σy, τxy] of a facing reaches the surface of the ellipsoid in Figure 6. failure surface
σ x
σ
y
τ
xy
Figure 6. Geometrical interpretation of the Tsai-Wu failure criterion.
PANEL COMPRESSION TESTS
A panel compression rig was built similar to a test frame for metal plates designed by A. C. Walker [2], see Figure 7. The rig is composed of a frame that supports the bottom and side edges of the panel, and a crosshead that slides in the frame, supports the top edge and loads the panel. In this way the crosshead is guided to prevent out-ofplane movements. Furthermore, top and bottom supports consist of sectioned slotted rollers supported by needle bearings and mounted in grooves in the base plate and the crosshead. The panel is subsequently inserted into the slots. The side edges are prevented from moving out-of-plane by knife-edge supports. Furthermore, the out-of-plane displacement at the panel centre is measured by a digital displacement gauge, see Figure 7 and Figure 8. Panels size 400x400 mm were cut from corrugated board and tested under compression with the cross-direction (CD) oriented in the direction of loading. Only flat panels with an imperfection less than half the thickness were selected for testing. Specimens were preconditioned for 24 o hours at 30% RH, 23 C, and subsequently conditioned for 24 hours at 50% RH, 23 o C, before testing. A total of 12 panels were tested and material and panel data is presented in Tables 1 and 2, respectively.
9
Figure 7. Panel compression rig and a corrugated board panel loaded to failure.
Figure 8. Measuring the deflection of a corrugated panel. Table 1. Material data for liner and fluting of the corrugated board.
Unit
Inner Facing Direction Kraft Liner Single Facer
Core SemiChemical Medium
Outer. Facing Kraft Liner Double Backer
Basis weight
g/m2
184.3
140.2
187.4
Thickness, t
mm
0.268
0.217
0.244
Elastic modulus, E 11
N/mm2
MD
7980
4750
8090
Elastic modulus, E 22
N/mm2
CD
3190
1560
2490
Tensile strength, σx,t
N/mm2
MD
81.4
46.9
82.1
Tensile strength, σy,t
N/mm2
CD
28.4
18.8
31.5
σx,c Compr. Strength,
N/mm2
MD
30.8
23.1
29.9
Compr. Strength, σy,c
N/mm2
CD
16.6
13.4
16.2
10
Table 2. Corrugated board data calculated using material data in Table 1. Transverse shear stiffness values are measured values using three-point bending [8]. Basis weight
g/m2
556
Thickness, h
mm
4.02
Corrugutation wavelength, λ
mm
7.26
Bending stiffness, D 11
Nm
14.6
Bending stiffness, D 22
Nm
5.43
Bending stiffness, D 12
Nm
2.71
Bending stiffness, D 66
Nm
3.34
Transverse shear stiffness, A 44
N/mm
39.2
Transverse shear stiffness, A 55
N/mm
5.6
RESULTS AND DISCUSSION
The load-displacement curves of the tested corrugated board panels show consistent buckling behaviour, see Figure 9. The dashed line is the analytical solution according to eq. (23) using the material and panel data in Tables 1 and 2, giving the critical buckling load P crit = 958 N, which is in accordance with classical buckling theory for orthotropic plates [4,11], and the post-buckling parameter Ψ= 8.6 N/mm2, see Table 3. Load [kN] 1.4
Predicted failure load 1.2
Experimental failure load 1.0
0.8
0.6
0.4
0.2
0 0
2
4
6
8
10
12
14
Out-of-Plane Displacement [mm]
Figure 9.
Load-displacement curves of 12 corrugated board panels. The dashed curve is the theoretical model with the analytical buckling load and postbuckling parameter. Failure predicted using the Tsai-Wu failure criterion. The dot-dashed curve is the theoretical model fitted to experimental curves.
11
Table 3. Analytically and experimentally determined parameters. Pcrit,exp
Pcrit
A0
Ψexp
Ψ
Pfail, exp
Pfail
N
N
mm
N/mm2
N/mm2
N
N
Average
814
958
0.8
3.55
8.6
1195
1265
Std
16
0.3
0.59
60
Max
844
1.3
4.68
1288
Min
786
0.5
2.88
1138
Unit
When the analytical out-of-plane deflection A has reached levels about half the plate thickness, the difference in analytical and experimental loads is about 20%. The brown dot-dashed line was determined by fitting the analytical expression in eq. (23), i.e. parameters P crit,exp , Ψexp and A0, to the experimental curves using non-linear regression. The regression was made using commercially available software called SAS [9]. The experimental critical buckling load P crit,exp = 814 N and the post-buckling coefficient Ψexp = 3.55 N/mm2 for the tested panels, see Table 3. The analytically and experimentally determined critical buckling loads differ by 18 %. The discrepancy between analytical and experimental post-buckling parameters in Table 3 is probably due to the non-linear response of the paper material at high stresses and local buckling of the facings, see Figure 10 [10]. However, the analytically calculated failure load P fail =1265 N differs only 6 % from the experimental failure load of P fail,exp = 1195 N. The analytical failure load was obtained by checking when the stresses [ σx, σy, τxy] in the inner facing satisfied the Tsai-Wu failure criterion in eq. (16). In Figure 11, failure of the inner facing is depicted by a range of colour fields indicating how close the material is to failure. This is expressed by the ratio between the length of the vector [ σx, σy, τxy] and an aligned vector that reaches the surface of the ellipsoid in Figure 6.
Figure 10. Local buckling of the facing on the concave side is visible just prior and after failure. 12
0
0 2
0 4
0 6
0 8
0 0 1
200 190 180 170 160 150 140 130 120 110 100 90 80 70 60 50 40 30 20 10 0 0 2 1
Edge
0 4 1
0 6 1
0 8 1
0 0 2
x
Centre
0,95-1 0,9-0,95 0,85-0,9 0,8-0,85 0,75-0,8 0,7-0,75 0,65-0,7 0,6-0,65 0,55-0,6 0,5-0,55 0,45-0,5 0,4-0,45 0,35-0,4 0,3-0,35 0,25-0,3 0,2-0,25 0,15-0,2 0,1-0,15 0,05-0,1 0-0,05
Figure 11. Colour fields indicate how close to material failure different areas of the inner facing are according to the Tsai-Wu criterion. Due to symmetry only the top-left quarter of the panel is shown.
CONCLUSIONS
A rig that furnishes simply supported boundary conditions has been designed to test corrugated board panels. Experimental results are consistent. An expression linking applied load with the out-of-plane deformation is derived. The first part of the expression is similar to ordinary Euler buckling of a column, where the maximum compressive load is limited by the critical buckling load. In the second part of the expression the membrane forces produce a parabolic relationship between the compressive load and out-of-plane displacement. The expression was fitted to experimental measured curves using non-linear regression to evaluate the critical buckling load and post-buckling coefficient for corrugated board panels. The results show an 18 % difference between experimentally estimated critical buckling load and the analytically predicted critical buckling load for orthotropic plates. This is partly attributed to excluded transverse shear deformation in the analytical solution. Compare the experimental value of 814 N with 870 N obtained from an analysis of a panel including transverse shear deformation [11,12]. A significant difference was also observed between analytically predicted and experimentally measured loaddisplacement curves at large out-of-plane deformation. This is probably caused by the non-linear material behaviour of paper and local buckling of the panel facings, i.e. the liners. However, the 6 % difference between the analytically calculated failure load and the experimental failure load is quite small. This suggests that collapse of the corrugated board panel is triggered by material failure of the inner facing. The strength of the material is therefore efficiently utilised. 13
REFERENCES
1.
M. Renman, "A mechanical characterization of creased zones of corrugated board", Licentiate thesis, Department of Engineering Logistics, Lund University, Sweden, 1994.
2.
A. C. Walker, "Thin Walled Structural Forms under Eccentric Compressive Load Actions" , Ph. D Thesis, University of Glasgow, June, 1964.
3.
J. Rhodes and J. M. Harvey, "Examination of Plate Post-Buckling Behaviour", Journal of the Eng. Mech. Div., vol. 103, No. EM3, pp. 461-478, 1977.
4.
R. M. Jones, "Mechanics of Composite Materials", Hemisphere Publishing Corp. New York, 1975.
5.
C. Fellers, B. S. Westerlind and A. de Ruvo, "An Investigation of the Biaxial Failure of Paper: Experimental Study and Theoretical Analysis", Transaction of the Symposium held at Cambridge: September 1981, Vol. 1, pp. 527-559, 1983.
6.
S. P. Timoshenko and J. M. Gere, "Theory of Elastic Stability" , McGraw-Hill, 1961.
7.
L. A. Carlsson, T. Nordstrand and B. Westerlind, "On the Elastic Stiffnesses of Corrugated Core Sandwich", Journal of Sandwich Structures and Materials, Vol. 3, pp. 253-267, 2001.
8.
T. Nordstrand and L.A. Carlsson, "Evaluation of Transverse Shear Stiffness of Structural Core Sandwich Plates", Composite Structures, Vol. 37, pp. 145-153, 1997.
9.
SAS Institute Inc., 100 SAS Campus DriveCary, NC 27513-2414, USA.
10.
P. Patel, T. Nordstrand, L. A. Carlsson, "Local buckling and collapse of corrugated board under biaxial stress", Composite Structures, Vol. 39, pp. 93110, 1997.
11.
T. Nordstrand, "On Buckling Loads for Edge Loaded Orthotropic Plates including Transverse Shear" , To be submitted to Composite Structures.
12.
T. Nordstrand, "Parametrical Study of the Post-buckling Strength of Structural Core Sandwich Panels", Composite Structures, Vol. 30, pp. 441-451, 1995.
14
APPENDIX
Marquerre's differential equation linking in-plane stresses with out-of-plane displacements for orthotropic plates is as follows 1 ∂ 4 F 1 + E *22 ∂ x 4 G12*
∂ 2 w = ∂ x∂ y
2
2ν 12* ∂ 4 F − * E 11 ∂ x 2∂ y 2
1 ∂ 4 F + = E 11* ∂ y 4
2 2 ∂ 2 w ∂ 2 w ∂ w0 − − 2 2 ∂ x ∂ y ∂ x∂ y
∂ 2 w0 ∂ 2 w0 − ∂ x 2 ∂ y 2
(1)
If the deflections w and w 0 are of the same form and their magnitudes are related by the expression w = A X ( x ) cos
π y b
(2)
and the initial imperfections as w0
π y b
= A0 X ( x ) cos
(3)
then a solution of eq. (1) for a plate with uniform compressive displacements of its ends, as shown in Fig. 4, is F = F 1 ( x ) + F 2 ( x )cos
2π y
b
(4)
Substituting eqs. (2-4) into (1) gives 2
1 π 2 II I 2 2 ( ) F = A − A X X + X 1 0 * 2 b E 22 1
IV
(
)[
2
IV
2
1 π 2 II I 2 2 A − A0 X X − ( X ) 2 b
where H =
)[
(
1 G 12*
−
2ν 12* E 11*
(5)
4
2π 2π 1 II F − HF + * F 2 2 2 * E 22 b b E 11 1
]
]
=
(6)
.
Thus F 1 is independent of y and hence constant along the length of the plate. While F 1 can not give a stress in the x-direction, it does give a stress in the y-direction which is found by integrating equation (5) twice.
A1
This gives
II
F 1
=
* E 22 π
2
( A
2 b
2
− A0
2
X 2
)
2
(7)
+ Bx + C
where B and C are constants and found from the membrane boundary conditions on the loaded ends. To obtain an expression for F 2 we assume that the deflections across the plate are in the form of a polynomial series N
X =
∑ A X n
(8)
n
n =1
where An
≤1
is the relative amplitude of the normalised shape function X n which in
turn is assumed to take the general algebraic form p
x X n = ∑ C pn a p 1 r
(9)
=
Substituting eq. (8) into eq. (6) gives 2
4
2π 2π 1 II F 2 − HF 2 + * F 2 * E 22 b b E 1
IV
=
11
2
1 π 2 2 A − A0 2 b
(
N N
)∑∑ A A ( X X n
m
n
II m
I
− X n
I
X m
)
(10)
n =1 m =1
Inserting the expression for X n , eq. (9), into eq. (10) and manipulating gives
2
4
2π 2π 1 II F 2 − HF 2 + * F 2 * E 22 b b E 1
IV
=
11
2
1 π 2 2 A − A0 2 b
(
)∑∑ An Am ∑ Bsmn x a n 1m 1 s 0 N N
=
=
2 ( r −1)
s
(11)
=
where r
Bsmn
=
∑ C(
s +2 − q ) n
Cqm[ q ( q − 1) − q ( s + 2 − q )]
q =1
A2
(12)
under the condition that C( s +2 −q <1) n
=
C ( s +2 −q >r ) n
=
0 . A solution for F 2 can be found by
putting F 2 in the same form as the right hand side of eq. (11), 2
F 2
1 π 2 2 A − A0 2 2a b
=
(
N N
)∑∑ A A φ n
m
mn
( x ) E 11* b 4
(13)
n =1 m =1
Substituting eq. (13) into eq. (11) gives 2
4
1 2π 2π 1 IV II φ mn − H φ mn + * φ mn * E 22 b b E 11
Bsmn x * 4 E 11 b a
2 ( r −1)
=
∑ s =0
s
(14)
If φ mn, part in turn is assumed to take the same form as the right-hand side of Eq. (14), 2 ( r −1)
φ mn, part =
x Lsmn ∑ a s 0
s
(15)
=
we obtain, by substituting eq. (15) into eq. (14) and equating coefficients from each term, for s = 2(r − 1) to 2( r − 1) − 1 Lsmn
=
Bsmn
( 2π)
(16)
4
and for s = 2(r − 2) to 2(r − 1) − 3 Bsmn H (s + 1)(s + 2 ) Lsmn
=
(2π )4
+
1 2π
2
(17)
L( s + 2 )mn
2 a E b * 11
and for 0 ≤ s ≤ 2(r − 1) Lsmn
=
Bsmn
+ 4
(2π )
H (s + 1)(s + 2) L( s + 2 )mn 2
−
1 2π
2 a E b * 11
1 −
* E 22
(s + 1)(s + 2)(s + 3)(s + 4) 1 2π
4
(18)
L( s + 4 )mn
4 a E b * 11
Eqs. (16) - (18) thus give the particular integral solution for eq. (14) when evaluated in the right order. To obtain a complete solution for φ , the complementary function solution must be added to eq. (15). There are three possible solutions to the homogeneous equation, 2
IV mn
φ
4
* 2π * 2π E 22 II − E 22 H φ mn + * φ mn b b E 11
consistent with the values of E 11* , E *22 and H . A3
=
0
(19)
* 2
2
1. If E 22 H
>
* 2 22
2
* 2 22
2
2. If E H
=
3. If E H
<
* E 22
4
, then all roots are real.
* E 11 * E 22
4
, then the roots are equal. This corresponds to the isotropic case.
* E 11
* E 22
4
, then all roots are complex.
* E 11
Only the first of these conditions is dealt with since this is consistent with the class of materials considered. Thus the solution to equation (19) is
φ mn,hom
= C 1mn
cosh k i x + C 2 mn sinh k 1 x + C 3mn cosh k 2 x + C 4 mn sinh k 2 x
(20)
where 1
* 2 2π 1 E 22 * * 2 2 k 1 = E 22 H + E 22 H − 4 * E 11 b 2 2
(21a)
and * 2π 1 E 22 * * 2 2 k 2 = E 22 H − E 22 H − 4 * E 11 b 2 2
1 2
(21b)
The complete solution for φ mn is 2 ( r −1)
φ mn
=
x Lsmn ∑ a s 0
s
+C 1mn
cosh k 1 x + C 2 mn sinh k 1 x +
=
+ C 3 mn
cosh k 2 x + C 4 mn sinh k 2 x
(22)
The values of the constants C imn are obtained by considering the boundary conditions at x=0 and x=a. It can be shown that if the unloaded edges are free to move in the plane of
the plate, the shear stress τ xy and the stress normal to the edge σ x must be 0.
∂ 2 F ∂ 2 F Since σ x = and τ xy = − , it follows by substituting eq. (13) into eq. (4) and ∂ y 2 ∂ x∂ y differentiate eq. (4) accordingly that I
I
φ mn (0 ) = 0; φ mn (a ) = 0; φ mn (0) = 0; φ mn (a ) = 0
(23)
The constants are thus found to be k1T3 (T1mn
)
+ T6mn − T4
C 1mn
=
C 2 mn
=−
C 3mn
= −C 1mn − L0 mn
(T2mn + T7 mn ) 2
T4 T5 − k1T3 T5 ( T1mn
)
+ T6mn − T3
(T2mn + T7 mn ) 2
T4 T5 − k1T3
A4
(24)
C 4 mn = −
k 1
C 1mn −
k 2
L1mn k 2 a
where 2 ( r −1)
T1mn =
∑L
smn
s =0
2 ( r −1)−1
T 2 mn =
∑ s =0
L( s+1)mn a
(s + 1)
T 3 = cosh (k 1a ) − cosh (k 2 a ) T 4 = sinh (k 1a ) −
k 1 k 2
sinh(k 2 a )
T 5 = k 1 sinh (k 1a ) − k 2 sinh (k 2 a ) T 6 mn = − L0 mn cosh (k 2 a ) −
L1mn k 2 b
T 7 mn = − L0 mn k 2 sinh (k 2 a ) −
sinh(k 2 a )
L1mn b
cosh(k 2 a )
The stress function F is completely determined in terms of the deflection given by eqs. (2) and (3). The constants in eq. (7) are found by compressing the plate uniformly along the loaded edges an amount v=-∆, see Fig. 4. During compression of the plate, it is assumed that no shear stresses are introduced along the loaded edges. b
b
Thus τ xy = 0 when y = − and y = and the compressive displacement must be 2 2 b
2 2 1 ∂ w 1 ∂ w0 + dy v = ∫ ε y − 2 2 ∂ y ∂ y b − 2
(25)
2
From Hooke's law we obtain 1 ν ε y = * σ y − 21* σ x E 22
(26)
E 22
and if
∂ 2 F ∂ 2 F ; σ y = σ x = ∂ y 2 ∂ x 2
(27)
a combination of equations (25) - (27) gives b
2 2 1 ∂ 2 F ∂ 2 F 1 ∂ w 1 ∂ w0 − + dy v = ∫ * −ν 21 2 2 2 2 E x ∂ y ∂ y ∂ y ∂ b 22 − 2
2
II
Substituting for F , w and w0 and integrating eq. (28), F 1 is found to be
A5
(28)
II
* v E 22 π
* 22
F 1 = E
+
b
2
2 2 2 ( A − A0 ) X 4 b
(29)
Substituting eq. (7) into eq. (29) gives v=
b
( Bx + C )
* E 22
(30)
To obtain uniform end displacement across the plate, B must be zero and the remaining coefficient will be * C = E 22
v
(31)
b
which is the stress induced in an unbuckled plate by the compression v. The strain energy in the buckled plate consists of two parts: the potential energy of bending and twisting, V B , and the membrane stretch energy, V M : b a 2
2 ∂ 2 ∂ 2 ∂ 2 1 D11 2 (w − w0 ) + 2 D11ν 21 2 (w − w0 ) 2 (w − w0 ) + V B = 2 0 b ∂ x ∂ x ∂ y
∫∫ −
2
∂ 2 + D22 2 ( w − w0 ) ∂ y b a 2
1 V M = ∫ ∫ * 2 0 b E − h
11
2
∂ 2 + 4 D66 ( w − w0 ) ∂ x∂y
∂ 2 F 2 ∂ y
2
2
dydx
(32)
2ν 12* ∂ 2 F ∂ 2 F + − * 2 2 E 11 ∂ y ∂ x
2
2
1 ∂ 2 F 1 + * + * G12 ∂ x∂ y E 22
∂ 2 F 2 ∂ x
2
dydx
(33)
Substituting the derivatives of F and w using using eqs. eqs. (2) - (4), (4), (8), (8), (13) (13) and and (29) (29) in in eqs. eqs. (32) - (33) and integrating with respect to y gives N b D11 a
2
V B = ( A − A0 )
N
∑ ∑ A A X 4 ∫ 0
n
m
II n
II
X m −
n =1 m =1
2
N N π II ν − 21 D11 ∑ ∑ An Am X n X m + b n =1 m =1
4 2 N N N N π π I I + D22 ∑ ∑ An Am X n X m + 4 D66 ∑ ∑ An Am X n X m dx b b n =1 m=1 n =1 m =1
A6
(34)
V M =
bh
( A 4
2
− A0
)
2 2
4
* N N N N π E 22 An Am A p Aq X n X m X p X q + ∫0 b 8 ∑∑∑∑ n=1 m =1 p =1 q=1 a
2
4 * 8 π E b N N N N II II + A A A A φ φ pq + * 4 ∑∑∑∑ n m p q mn b 4 E 22 a n=1 m=1 p =1 q=1 11
4
4
*
2π π E 11b + 4 b b 4a 2
4
*
8 N N N N
∑∑∑∑ A A A A φ n
p
q
φ +
mn pq
n =1 m =1 p =1 q=1
* 8 N N N N 11
2π π ν E b 4 b b 2a
+
m
12
∑∑∑∑ A A A A φ n
m
p
q
II mn pq
φ
+
n =1 m =1 p =1 q =1
2 4 2π π E b 8 N N N N I I + A A A A φ φ pq dx + 4 ∑∑∑∑ n m p q mn b 4 b G a n =1 m =1 p =1 q=1 *2 11 * 12
v + 2E b
2 a N N 2 2 π bh * v An Am X n X m dy ∑ ∫0 dx + ( A − A0 ) b 4 E 22 b ∫0 ∑ n m 1 1 = =
2a
* 22
(35)
The total strain energy is obtained by adding V B and V M , V = V B + V M
(36)
In order to minimise this expression with respect to A we have to know the values of Ai , i=1..N, i.e. the relative amplitudes of the shape functions. Assuming that the imperfect plate will have the same deflected shape as a perfect plate, with A0 = 0 , we can obtain an accurate description of Ai by the first buckling mode. Initially, limited bending can be produced with negligible membrane stretching of the middle plane, and we need to consider only the bending energy and the corresponding work done by the external forces acting in the middle plane of the plate, i.e. all terms that include v in eq. (35). Subsequently, all terms in eq. (36) that contain the fourth power of Ai can be omitted. Thus equation (36) can be written as 2 a π II II II V = An Am ∫ [ D11 X n X m − 2 ν 21 D11 X n X m + ∑ ∑ 4 n=1 m=1 b 0 2 4 2 2 * E 22 π π π * v v I I + 4 D66 X n X m + D22 X n X m + E 22 h X n X m dy + abh (37) 2 b b b b b
bA2
N N
The minimum of V is found by differentiating V with respect to Ai and using dV =
∂V ∂ A1
dA1 +
∂V ∂ A2
dA2 +. .. . . .+
∂V ∂ An
dAn = 0
(38)
which in turn gives N linear simultaneous equations, from which the value of v = −∆ crit to cause buckling can be found. The relative values of Ai can also be obtained. Thus
A7
2 2 4 π π π II II II I I S nm = ∫ D11 X n X m − ν 21 D11 X n X m + 4 D66 X n X m + D22 X n X m dx b b b 0 a
(39)
π 2 * h = ∫ E 22 X n X m dx b 0 b a
Qnm
gives
S + S T − 2∆ Q A = 0 crit
(40)
and the non-trivial solution of eq. (42) demands that T
S + S − 2∆ crit Q = 0
(41)
Solving eq. (41) for ∆ crit and inserting the value of ∆ crit in eq. (40) enable us to obtain the normalised eigenvector A . Thus all Ai :s are determined. Inserting the values of Ai in equation (36) and minimising V with respect to A yields the relationship between A and v as follows:
A0 C 2 A C 1
v = −1 −
(
2
− A − A0
2
)C
3
(42)
C 1
where 2
a * h N N π E 22 C 1 = An Am ∫ X n X m dy ∑ ∑ 8 b n =1 m =1 0
(43)
2 a π II II II C 2 = ∑ ∑ An Am ∫ D11 X n X m − ν 21 D11 X n X m + 8 n=1 m=1 b 0 2 4 π π I I + 4 D66 X n X m + D22 X n X m dy b b
b
N N
(44)
4 * a E 22 π bh N N N N C 3 = An Am A p Aq ∫ X n X m X p X q + ∑∑∑∑ b 4 8 n =1 m=1 p=1 q =1 0
+
*2 8 E 11 b 1
* 4a 4 E 22
″
φ mn φ pq
″
2π + b
2
2ν 12* 1 ″ ′ ′ E * φ mnφ pq + G * φ mn φ pq + 11 12
2 2π 1 + * φ mnφ pq dx b E11
(45)
A8
Paper 6
_______
Local Buckling and Collapse of Corrugated Board under Biaxial Stress by P. Patel, Nordstrand, T. and Carlsson, L.A. Composite Structures, Vol. 39, No. 1-2, pp. 93-110, 1997.
Detta är en tom sida!
Paper 7
_______
A Strength Prediction Method for Corrugated Board Containers by T. Nordstrand, Blackenfeldt, M. and Renman, M. Report TVSM-3065, Div. of Structural Mechanics, Lund University, Sweden, 2003.
Detta är en tom sida!
A Strength Prediction Method for Corrugated Board Containers Tomas Nordstrand, Mikael Blackenfeldt and Magnus Renman SCA Research, Box 716, 851 21 Sundsvall, Sweden
ABSTRACT
A method for prediction of the top-to-bottom compression strength of corrugated board containers using finite element analysis was developed. Up to triple-wall corrugated board is accommodated in the finite element model. In order to keep computational time down, the corrugated core layer(s) is considered to be homogeneous with effective material properties, and the complete corrugated board is represented by a multi-ply eight node isoparametric shell element. The shell elements represent the side panels of the box. Coupling elements are used to represent creases at the top and bottom edges. Buckling and large displacements are considered in the analysis, and container collapse loads are predicted using the Tsai-Wu failure criterion. It is assumed that failure is triggered by material failure in one of the facings of the corrugated board. Local buckling of the corrugated board constituents is not considered in this study, and orthotropic linear elastic material properties are assumed for each layer. The boxes analysed fail in a post-buckled state. A non-linear finite element method involving Newton-Raphson iterations is discussed in conjunction with the buckling analysis. Convergence of the predicted collapse load with the number of elements used in the model is studied, as well as the sensitivity of the collapse load to the imposed stiffness at the loaded boundaries. Different buckling modes of a box are simulated, giving an in-depth understanding of the relation between the strength of a box and constraints imposed on the panels by the corners of the box. Finally, the results of an extensive testing program comprising about 1,300 box strength tests are summarised and used to verify the container strength prediction method developed.
1
INTRODUCTION
Although corrugated board has been used for over a century, the packaging design process is still mostly empirical or semi-empirical. Research on corrugated board has been lagging behind its industrial application [1]. There are many reasons for this. One is that the mechanics of corrugated board and packaging are quite complex due to the structure of the board, difficulties in loading and control of boundary conditions. Consideration of the corrugated board market in recent years, with fewer companies increasing their market shares, has been a driving force for improvement of design procedures. Production volumes have increased considerably. Even small savings in raw material achieved through improved design procedures imply significant cost savings [2]. In addition, with the development of powerful computing tools and structural analysis codes such as finite element software, it is now possible to obtain numerical solutions for the stresses and deformations in corrugated board structures when loaded [3]. The benefits of such an analysis are apparent since the analysis is based on the geometry of the corrugated board and the physical properties of the constituent liners and corrugated medium. In contrast, the empirical approach to design requires manufacture of the board and tedious making and testing of boxes and specimens, cut from board [4]. This work was initiated with the objective of developing a design method based on finite element analysis to predict top-to-bottom compression strength of corrugated containers. The finite element model includes multi-ply eight node isoparametric shell elements for the side panels and coupling elements representing creases at top and bottom edges [5, 6]. Creases are scored folding lines between flaps and sides. McKee and Gander [4] found that 90% of the compression deformation of a box occurs in the creases. The residual compressive and rotational stiffness of the creases, being most important [7], are accommodated using coupling elements [5]. Extensional, shear and bending stiffnesses of the corrugated board are calculated from the geometry and material properties of the constituent liners and medium [8-11]. The material properties are assumed to be orthotropic linear elastic. In a previous study of collapse of corrugated board panels it was shown that the TsaiWu failure criterion [12-13] could be used to predict material failure in one of the facings of the corrugated board panels. The same failure criterion is used in the present analysis of boxes, and local buckling of the outer facings is disregarded [14]. The buckling of side panels requires a non-linear finite element analysis that is solved using the Newton-Raphson method [15, 16]. The solution is terminated when the failure criterion is satisfied. Convergence of the predicted collapse load with the number of elements used in the model is studied, as well as the sensitivity of the collapse load to the imposed stiffness at the loaded boundaries. Different buckling modes of a box are simulated, giving an in-depth understanding of the relation between the strength of a box and constraints imposed on the panels by the corners of the box.
2
Finally, extensive testing of boxes made from two different corrugated board grades is used for comparison with finite element predictions. A commercially available finite element code called ANSYS [5] was chosen to solve these tasks for two main reasons. The software is adapted to operate in Microsoft’s Windows NT environment, and the ANSYS Parametric Design Language (APDL) [5] makes it easy to implement algorithms for processing the finite element models.
NUMERICAL MODEL Material stiffness properties
Up to triple-wall corrugated board is accommodated in the finite element model. To keep computational time down the corrugated core(s) is homogenised [8-11] and the complete corrugated board is represented by a multi-ply eight node isoparametric shell element, SHELL99 [5]. Each layer, from three to seven, of the homogenised corrugated board is described by its thickness and orthotropic elastic properties, i.e. the elastic moduli (EX, EY, EZ), Poisson’s ratio (NUXY, NUYZ, NUXZ) and the shear moduli (GXY, GYZ, GXZ). The elastic planes of symmetry are aligned with the machine direction(MD) and cross direction (CD) of the board, see Figure 1.
ZD
Z y
CD hc
λ
X
MD
Figure 1. Homogenisation of the corrugated core of single-wall corrugated board.
Model and boundary conditions used for a regular slotted container
In a finite element analysis of a corrugated board container, a quarter of the box is represented by i xk+jxk 8-node elements, i=j and k=2i, where i represents the number of elements parallell with the x-axis, j the number of elements parallell with the y-axis and k the number of elements parallell with the z-axis, see Figure 3. Then the number of nodes is (16+5(i+j-2)+10(k-1)+3(i+j-2)(k-1)). Each node has six degrees of freedom, viz. translation and rotation along axis parallell to the nodal coordinate system. When considering a box loaded under compression, it is recognised that the creases at the top and bottom affect the loading of the box [4]. In order to predict topto-bottom compression of boxes the properties of the creases were measured in earlier work [7]. The soft creases are accommodated in the FEM-model by edge springs attached at the loaded top and bottom edges. The springs are modelled using element COMBIN14 [5] and the stiffness of the springs is adjusted in order to introduce consistent nodal forces on the shell elements if uniformly compressed [16]. The creases also introduce an eccentricity of the applied loads [7], which is simulated by 3
introducing a moment that is constant at the top and bottom edges using element COMBIN37 [5], see Figure 2. The vertical edges at the four corners of the box are connected in translation only with no rotational constraint.
Figure 2. Elements used to simulate the response of the creases. A quarter of the box is analysed due to symmetry. Boundary conditions are imposed on the nodes along the edges and nodes at the planes of symmetry according to Figure 3. Only the two symmetry planes indicated in the figure are utilized since the boxes analysed in general have more complex geometry than the box shown in figure 3.
k=16
=8
i=8
Figure 3. Quarter of a box model with boundary conditions: L9: DX=0, DZ=displacement step, (nodes are free to move in y-direction except where L9 L12: DY=0, DZ=displacement step L2: DX=0, DZ=0 L3: DY=0, DZ=0 L10: Translation coupling of the nodes at the corner . L7: DY=0, RotX=0, RotZ=0 (symmetry) L13: DX=0, RotY=0, RotZ=0 (symmetry) L9, L12, L2, L3: Connected to elements shown in Figure 2.
4
intersects L7)
Analysis
The stress-strain relationship of each layer, j, can be expressed by [8]
{σ } j = [Q] j {ε } j
(1)
where {σ} j and {ε} j are the vector stress and strain, respectively, in a plate. Plane stress is assumed and the elastic stiffness matrix [Q] j is defined as
E x, j ν E yx, j x, j 0 1 [Q] j = 1 - ν xy, jν yx, j 0 0
ν xy, j E y, j
0
0
E y, j
0
0
0
G xy, j
0 G yz, j
0
0
0
0
f 0
0 0
0
0 G xz, j f
(2)
where ν xy,j and ν yx,j are the Poisson's ratios, E x, j and E y, j are the elastic moduli and Gxy, j , Gyz, j and Gxz, j are the in-plane and transverse shear moduli, respectively. In a
finite element analysis where the transverse shear strain is assumed to be constant through the thickness of each layer, it is common to reduce the transverse shear moduli by a "shear correction factor" of 1.2 to compensate for the excessive amount of shear strain energy produced [5]. In the current finite element formulation the transverse shear moduli of each layer are reduced by a factor f given by A f = 1.2 or f = 1.0 + 0.2 whichever is greater, where A is the area of the element 25h 2 and h the thickness of the panel. The latter expression is included in order to prevent shear locking [5]. The homogeneous properties of the corrugated core layers are obtained according to ref. [8]. The stiffness matrix can be integrated through the thickness h of the panel to o btain the extensional, coupling and bending stiffness matrices as follows N
[A ] = ∑ [Q] j (z j − z j-1 )
(3a)
j=1
1 N [C] = [Q] j (z j2 − z j2-1 ) 2 j=1
(3b)
1 N [D] = [Q] j (z j3 − z j3-1 ) 3 j=1
(3c)
∑
∑
where z j are the ply coordinates [8] and N= 3, 5, 7. These stiffness matrices in turn can be combined with the strain-displacement matrix [B], which connects the displacements to the strains and curvatures of the element, to form the element stiffness matrix [k] as follows
5
[k ] = ∫ ([B0 ]T [A][B0 ] + [B0 ]T [C][B1] +[ B1]T[ C][ B0] +[ B1] T[ D][ B1] ) dA
(4)
A
where [B0]+z[B1] = [B], see e.g. [16]. The element stiffness matrices are subsequently assembled to a global stiffness matrix [K], which is used in an eigenvalue analysis of the finite element model to determine a suitable load step for the non-linear analysis. The following equation is solved in the eigenvalue analysis
([K ] + χ [S]ref ){ψ } = {0} where S
ref
(5)
is obtained by applying consistent nodal forces {P}ref on the top and bottom
edges corresponding to a uniform unit pressure, and performing a static linear analysis to obtain the membrane stresses that generate the "stress stiffness matrix" S [16]. χ ref
is the factor used to multiply the loads that generate the stresses, and {ψ } is the generalised displacement vector for the nodes [16]. The load {P} is scaled by the factor χ which also alters the intensity of the membrane stresses but not the distribution of the stresses such that
{P} = {P}ref ⇔ [S] = χ [S]ref
(6)
As χ is increased, the overall stiffness of the box, ([Κ] + [S]), is reduced until a critical load {P}cr corresponding to the eigenvalue χcr is reached and the box becomes unstable, i.e. det ([Κ] + [S]) goes to zero. The panels collapse at a load much higher than the critical buckling load P cr. In this socalled postbuckled state, the deflection of a panel is large enough to introduce geometric non-linearities between the load and displacement of the panel. This means that the stiffness matrix [Κ] of the box becomes a function of the unknown displacements, i.e.
[K (ψ )]{ψ } = {P}
(7)
where {ψ } is a set of known and unknown nodal displacements and {P} represents a set of external loads that act on the nodes. In order for the finite element model of the box to exhibit out-of-plane deformation, it must contain an imperfection, which in our case is introduced as an eccentricity of the applied loads [7] simulated by introducing a moment that is proportional to the load at the top and bottom edges using element COMBIN37. The nonlinear problem is solved step-wise using the Newton-Raphson method [15, 16]. The general algorithm proceeds as follows [5]. 1. Prescribe known displacements {ψ k }i of the nodes on the edges parallel to the yaxis, in this case a uniform displacement. The superscript k denotes known displacements and the subscript i is the load step. 2. Compute the unknown forces {P}i and displacements {ψ u }i using the stiffness matrix from the earlier step K i-1 (For the first step, i = 1, the global stiffness matrix of the un-deformed panel is used K 0 ).
6
[K ]i-1 ({ψ k }i + {ψ u }i ) = {P}i 3. Compute the updated stiffness matrix K i corresponding to the configuration {ψ k }i + {ψ u }i . 4. Calculate restoring force {P r }i and the change in displacements {∆ψ }i as follows [K ]i ({ψ k }i + {ψ u }i + {∆ψ }i ) = {P}i + {P r }i . 5. Convergence is checked by Max {∆ψ }i ≤ c where c is an arbitrary small number for which equilibrium is approximately obtained. 6. If the solution is not converged, add {∆ψ }i to {ψ u }i , add {P r }i to {P}i and repeat steps 3 - 5. 7. If the solution is converged continue with the next load step {ψ k }i +1 = {ψ k }i + {ψ u }i + {∆ψ }i + {∆ψ k }i+1 . The loading is terminated when the stress-state at any point in the facings of the panels exceeds a predefined failure criterion. Failure criterion
In this study the Tsai-Wu tensor failure theory [13] is used to obtain the ultimate loadcarrying capacity of a box. According to this theory, in the plane stress condition, material failure develops when f=1, where (8)
2 f = Γ 1σ x + Γ 2σ y + Γ 11σ x2 + Γ 22σ y2 + Γ 66τ xy + 2Γ 12σ xσ y
where Γ 11 = −
Γ 2 =
1 σ y,t
+
1 σ x,tσ x,c
1 σ y,c
, Γ 66 =
, Γ 22 = −
1 σ y,tσ y,c
, Γ 12 = − 0.36 F11F22 , Γ 1 =
1 σ x,t
+
1
,
σ x,c
1 σ x,cσ y,c
The subscript j=t or j=c of strength σi,j ,i=x,y, denotes the strength in tension and compression, respectively. The expressions for Γ 12 and Γ 66 are approximations for paper materials [13]. It is assumed that the box fails as soon as this failure criterion is fulfilled in any point of the outer facings of the corrugated board.
EVALUATION OF MODEL
Boxes of different sizes and board grades, B-140T/112R/140T, C-140T/112R/140T, C140K/112R/140K and C-180L/112R/180L were studied. The corrugated board is coded, the first letter indicates the geometry of the core, see Table 2. The three following sets of a number and a letter indicate the liner and fluting composition of the board, see Table 1. Material data for the liners and fluting, measured according to the Scandinavian pulp, paper and board standard SCAN-P 67:93 and SCAN-P 46:83, is shown in Table 1 and geometry of the corrugated cores in Table 2. Observe that the compression strength is assumed to be 0.78*SCT (Short-span Compression Test) due to the small specimen size used in the SCT [17]. The finite element models do not
7
include any imperfections. Instead, eccentricity moments are introduced at the top and bottom edges. Evaluation of the element density requirements
When performing a finite element analysis the chosen number of elements is a balance between accuracy and calculation time. A larger number of elements could give a more precise solution, but the calculation time increases dramatically with the number of elements. In order to check the influence of element size on the calculated failure load, finite element solutions were generated for boxes with a size of 300 x300x300 mm (LengthxWidthxHeight). Eccentricity moments bending the panels outward, obtained in Ref. [7] as the residual after folding the flaps, are introduced at the top and bottom edges with 0.6 Nmm/mm for B-140T/112R/140T and C-180L/112R/180L. Stiffness of the creases is 20.0 (N/mm)/mm for both boards. As can be seen in Table 3, the number of elements has little influence on the box strength estimates. The largest difference is for the 16 element B-140T/112R/140T box, suggesting that the element size is more important for light and thin board grades than for heavy and thick board grades. Figure 4 shows the effect of element size on the Tsai-Wu failure criterion f , eq.(16), where 16 elements give too low precision. So generally, 16 elements seem to be too few but 64 elements and upwards give good accuracy for both B- and C-flute. Table 1. Material data for liners and flutings (Compression strength is 0.78*SCT and stiffness in compression is set equal to stiffness in tension) 140 T Testliner 3 2
Grammage (g/m ) Thickness (µm) Tensile stiff MD (kN/m) Tensile stiff CD (kN/m) Tensile str MD (kN/m) Tensile str CD (kN/m) SCT MD (kN/m) SCT CD (kN/m)
140 K Kraftliner 2
Grammage (g/m ) Thickness (µm) Tensile stiff MD (kN/m) Tensile stiff CD (kN/m) Tensile str MD (kN/m) Tensile str CD (kN/m) SCT MD (kN/m) SCT CD (kN/m)
Avg 136 225 830 460 6.4 3.8 3.3 2.4
Sd 0.9 8.1 18 10 0.71 0.43 0.22 0.18
n 30 30 10 10 10 10 10 10
180 L Testliner 2
Avg 139 195 1390 490 14.0 5.3 5.1 2.7
Sd 1.2 5.8 23 21 0.31 0.24 0.12 0.08
n 30 30 10 10 10 10 10 10
112 R Recycled fluting
2
Grammage (g/m ) Thickness (µm) Tensile stiff MD (kN/m) Tensile stiff CD (kN/m) Tensile str MD (kN/m) Tensile str CD (kN/m) SCT MD (kN/m) SCT CD (kN/m)
2
Grammage (g/m ) Thickness (µm) Tensile stiff MD (kN/m) Tensile stiff CD (kN/m) Tensile str MD (kN/m) Tensile str CD (kN/m) SCT MD (kN/m) SCT CD (kN/m)
8
Avg 178 283 1190 550 9.97 4.63 4.98 3.29
Sd 1.3 9.5 18 15 0.98 0.54 0.21 0.16
n 30 30 10 10 10 10 10 10
Avg 114 199 911 354 7.32 2.9 3.8 2.12
Sd 0.87 8.2 19 13 0.26 0.16 0.16 0.12
n 30 30 10 10 10 10 10 10
Table 2. Flute profiles. α is the ratio of flute to liner length. Profile Wavelength, λ (mm) Flute height, h c (mm) Take-up factor, α
B 6.3 2.46 1.32
C 7.8 3.61 1.43
Table 3. Failure load (N) of the whole box versus number of elements. 300x300x300 mm, C-180L/112R/180L No. of elements FEM (N) Index 16 4768 1.04 36 4631 1.01 64 4630 1.01 144 4605 1.00 256 4590 1.00 300x300x300 mm, B-140T/112R/140T No. of elements FEM (N) Index 16 2516 1.15 36 2275 1.04 64 2188 1.00 144 2181 1.00 256 2165 0.99
Figure 4. Tsai-Wu value f of the inner facings, eq.(16), for the 300x300x300 mm box, B-140T/112R/140T with 16 and 144 elements respectively.
Evaluation of Tsai-Wu failure criterion vs. compression load
The collapse load of boxes size 300 x300x100 mm and 300x300x300 mm is predicted using the Tsai-Wu failure criterion, eq. (8). Eccentricity moments bending the panels outward are introduced at the top and bottom edges with 0.6 Nmm/mm for C-
9
140T/112R/140T. Crease stiffness is 20 (N/mm)/mm. In order to make the analyses converge within reasonable time, a Tsai-Wu value between f = 0.95 and f = 1.05 is used. Typically, the maximum difference in strength predictions between f = 0.95 and f = 1.05 is within 4%. However, when small differences between subsequent predictions need to be studied a more accurate prediction is necessary. Figures 5 and 6 show typical load vs. Tsai-Wu curves and the load vs. out-of-plane displacements. Since there is almost a linear relationship between the Tsai-Wu value and the load in Figures 5 and 6, the result from the FEM-calculation, P i , can be adjusted using the present and previous Tsai-Wu value, f i and f i-1 and load step ∆P to obtain the failure load. P f =
∆P (1 − f i ) + Pi f i − f i −1
(9)
The box strength estimated using eq. (9) is P f = 3052 N for C-140T/112R/140T, 300x300x300mm and P f = 4873 N for C-140T/112R/140T, 300x300x100mm. For the boxes in chapter 4 ( Appendix Table A1 and A2) the maximum correction with eq. (9) is 2.9% and the average correction is 1.2%.
out-of-plane displacement panel center (mm) 0
350
1
2
3
4
5
6
7
8
9
300 250
) N ( P 200 , d a o L 150
Tsai-Wu out-of-plane
100 500 0 0
0.2
0.4
0.6
0.8
1
Tsai-Wu, f
Figure 5. Load vs. Tsai-Wu value f and out-of-plane displacement for C140T/112R/140T, 300x300x300mm. P f = 3052 N. Since eccentricity moments bending the panels outward are introduced at the top and bottom edges, the Tsai-Wu value and out-of-plane displacement are not zero at zero load.
10
A out-of-plane displac ement panel center (mm) 0
500
0.1
0.2
0.3
0.4
0.5
0.6
0.7
450 400 350
) N ( P 300 , d a 250 o L
Tsai-Wu out-of-plane
200 150 100 500 0 0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1
Tsai-Wu, f
Figure 6. Load vs. Tsai-Wu value f and out-of-plane displacement for C140T/112R/140T, 300x300x100mm. Pf = 4930 N.
Additional failure criterion
If boxes do not buckle, failure load is limited by the crushing strength of the corrugated board, see Figure 7. The crushing strength of the creased corrugated board is about 50% of the edge crush test strength (ECT) [7], and the ECT can be predicted from the short-span compression test (SCT) of the board constituents [17]. Thus the crushing strength per unit length of the creased corrugated board, CBS , in the CD, see Figure 7, is M 0.85 0.78SCT1 + ∑ 0.78(α 2k SCT2 k + SCT2 k +1 ) CBS = 2 k =1
(10)
where α2k is the take-up factor of the corrugated core layer 2k and when M = 1, 2, 3 it corresponds to single, double or triple-wall board, see Figure 1. For board C140T/112R/140T CBS = 2.59 N/mm. In chapter 3.2 the crushing strength of box no.1 C-140T/112R/140T, 300x300x100mm, see Appendix Table A2, is P crush = 3108 N and corresponding experimental value is P f,exp = 2932 N. Pcrush = 3823 N and P f,exp = 3290 N for box no.13 C-180L/112R/180L, 300 x300x100mm, see Appendix Table A2.
Figure 7. A crushed shallow box.
11
Influence of the edge stiffness
The edge/crease stiffness is an important property for box compression strength because it affects the load distribution on the top and bottom edges when the side panels buckle. Thus, a soft edge prevents a redistribution of the load towards the corners of a box during buckling. This increases the bending of the panels and reduces box strength. Edge/crease stiffness can be measured by using a special rig [7]. Figures 8a and 8b show the predicted failure load as a function of the edge stiffness for a box with B- and C-board, respectively. The predicted failure load increases rapidly with stiffness until stiffness gets close to the stiffness of a cut-edge. For the B-flute box the strength increases by about 16% between stiffness 2 and 100, whereas for C-flute the increase is about 8%. 3150
3100
10 8
) N ( 3050 , f P ) N d ( a M3000 o l E F e r u l i 2950 a F
10 6
2900 10 0 2850 0
10
20
30
40
50
60
70
80
90
100
Edge stiffness (N/mm² )
Figure 8a. Predicted failure load for B-140T/112R/140T, 300x300x300 mm. An index is included for stiffness 2.0, 20.0 and 100 (N/mm)/mm. 2300 2250
116
) N ( 2200 , f P ) d 2150 N a ( o l M E e r F 2100 u l i a F 2050
112
2000 10 0
1950 0
10
20
30
40
50
60
70
80
90
100
Edge stiffness (N/mm²)
Figure 8b. Predicted failure load for C-140T/112R/140T, 300x300x300 mm. An failure index is included for stiffness 2, 20 and 100 (N/mm)/mm.
12
Influence of eccentricity moments orientation on buckling modes
Applying different eccentricity moments at the loaded edges triggers different buckling modes. Two alternative moment directions are used corresponding to a panel buckling direction inwards or outwards, see Figure 9. Results in Table 4 show that outward moment for all panels clearly gives the strongest box and inward moment gives the weakest box. Since the panels are completely identical and the value of the moment is 0.6 Nmm/mm in all cases, the difference in strength is only caused by geometrical differences in the buckling patterns. Displacements of the box corners in particular will be different, see Figure 10.
in
out
Figure 9. Applied eccentricity moments. Table 4. Boxes size 300x300x300 mm and symmetrical board C-140K/112R/140K Driving moment all out all in 2 out/2 in 3 out/1 in 3 in/1 out
Failure load, N 3942.9 3565.8 3673.5 3671.8 3566.3
13
Index 100 90 93 93 90
Buckling all out all in 2 out/2 in 3 out/1 in 3 in/1 out
Displacement of the corners (mm) -0.21 -0.16
7.2
-0.18
-0.22
-7.9
Out-of-plane displacement at panel centre (mm)
7.0
-6.9
Figure 10. Schematics of deformed corner cross-sections at the centre of the box. The dot-dashed lines indicate original position of the corners. The schematics are not drawn to scale. The three first buckling modes in Table 4 are shown at index load 90. If the ratio between height and width of the panels is different panels can buckle in completely different ways, see Figure 11.
1/4 model 300x300x300
1/4 model 400x200x300
Figure 11. Tsai-Wu value f of the inner facings, eq.(16). Different buckling modes of boxes with the same perimeter.
14
EXPERIMENTS AND VERIFICATION OF FINITE ELEMENT MODEL
Boxes made from board with either a B-flute or a C-flute profile, see Table 2, have been tested to verify the finite element model. About fifty different box geometries for each flute profile and a total of 1,316 boxes were tested. All boxes were unprinted and manually creased, slotted, glued and erected. All boxes were conditioned at 23°C and at 50% RH for at least 24 hours. Some of the tests were performed at SCA Packaging Munksund with a rigid platen box compression tester, and some at SCA Packaging Research, Aylesford, with a floating platen (rotational unrestrained) box compression tester. Compression speed was 10 mm per minute and time to failure was about one minute. Specifications, box compression test (BCT) results and FEM estimates of the tested boxes are listed in Table A1 and Table A2 in the Appendix together with material data for the liners and mediums in Table A3. All predicted strengths are based on an edge stiffness of 2 N/mm 2 and an element density of 144 elements. Moment applied at the loaded edges is 0.6 Nmm/mm [7] in the conservative configuration in/out, see Table 4. Figures 12 and 13 show comparisons between box strength estimated with FEM and the strength of tested handmade boxes according to Table A1 and A2, respectively. On average, the finite element simulation predicted strength 3 % above the tested average of boxes in Figure 12 and 5 % below average of boxes in Figure 13. However, two of the tested boxes in Figure 13 (box no.1 and 13) have a strength that is significantly less than predicted. These boxes are shallow, which makes them less inclined to buckle and they fail due to crushing instead, see chapter 3.3. No explanation has yet been found for the over-conservative predictions for box geometries Nos. 29, 30 and 49.
5000 ) N ( 4500 h t g n 4000 e r t S n 3500 o i s s e r 3000 p m o C 2500 x o B 2000
Test FEM
1500 1000 0
10
20
30
40
50
Box geometry no.
Figure 12. FEM-estimate box strength compared to test results for a range of handmade boxes according to Table A1. Flute profile B, see Table 2.
15
7000 6000 ) N ( 5000 h t g n e r 4000 t S n o i 3000 s s e r p m 2000 o C
Test FEM
1000 0
0
5
10
15
20
25
30
35
40
45
50
Box geometry no.
Figure 13. FEM-estimate box strength compared to test results for a range of handmade boxes according to Table A2. Flute profile C, see Table 2.
CONCLUSIONS AND REMARKS
A finite element model of corrugated board containers is shown to predict the failure load of boxes, made from B- and C-board, within an average error margin of 5%. Effective material properties of the homogenised corrugated cores have been used, and each layer of the corrugated board is assumed to be orthotropic linear elastic. It is shown that convergence is obtained with relatively few elements, e.g. 64 elements are quite sufficient for a regular size box, i.e. 300 x300x300 mm. The edge stiffness has a significant influence on the predicted failure loads because it affects the load distribution on the top and bottom edges when the side panels buckle. There is also a variation of about 10% in the failure load due to different buckling modes, see Table 3. This is attributed to different constraints imposed on the side panels by the corners of a box. Boxes of different sizes and board grades were tested and compared to predicted strengths. On average, the difference between experimental and predicted values was small. However, for some boxes the difference was significant because the boxes did not buckle but were crushed instead. The described FE-method for predicting the top-to-bottom compressive strength of corrugated containers has been used as the basic component in the subsequent development of a user-friendly computer-based tool for strength design of containers. The graphical interface of this design tool is shown in Figure 14. The example shows menus that relate to input data definition.
16
Figure 14. The graphical interface of a design tool for corrugated boxes.
REFERENCES
1.
R. Steadman, "Corrugated Board", Ch. 11, in Handbook of Physical Testing of Paper, (R.E. Mark et al. eds.), pp. 563-660, Marcel Dekker, New York, 2002.
2.
B. Steenberg, J. Kubat, S. Martin-Löf and L. Rudström, "Competition in Rigid Packaging materials", Svensk Papperstidning., Vol. 73, No. 4, pp. 77-92, 1970.
3.
T. Nordstrand, "Parametrical Study of the Post-buckling Strength of Structural Core Sandwich Panels", Composite Structures, No. 30, pp. 441-451, 1995.
4.
R. C. McKee and J. W. Gander, "Top-Load Compression", TAPPI, Vol. 40, No. 1, pp. 57-64, 1957.
5.
ANSYS User's Manual, Swanson Analysis Systems Inc., Vol.1-2, 1989.
6.
O. C. Zienkiewicz, " The Finite Element Method in Engineering Science", McGraw-Hill, 1971.
7.
M. Renman, "A mechanical characterization of creased zones of corrugated board", Licentiate thesis, Department of Engineering Logistics, Lund University, Sweden, 1994.
17
8.
L. A. Carlsson, T. Nordstrand and B. Westerlind, Westerlind, "On the Elastic Stiffnesses Stiffnesses of Corrugated Core Sandwich", Journal of Sandwich Structures and Materials, Vol. 3, pp. 253-267, Oct., 2001.
9.
C. Libove and R. E. Hubka, "Elastic "Elastic Constants for Corrugated Corrugated Core Sandwich Plates", NACA TN 2289, 1951.
10. R. Paetow and L. Göttsching , " Poisson's Ratio of Paper", Das Papier a. (in German), Vol. 6, pp. 229-237, 1990. 11. T. Nordstrand and L.A. Carlsson, "Evaluation of Transverse Shear Stiffness of Structural Core Sandwich Plates", Composite Structures, Vol. 37, pp. 145-153, 1997. 12. T. Nordstrand, "Analysis and Testing of Corrugated Corrugated Board Panels into the Postbuckling Regime", To be submitted to Composite Structures . 13. C. Fellers, B.S. Westerlind and A. de Ruvo, "An Investigation of the Biaxial Failure of Paper: Experimental Study and Theoretical Analysis", Transaction of the Symposium held at Cambridge: September 1981, Vol. 1, pp 527-559, 1983. 14. P. Patel, T. Nordstrand and L.A. Carlsson, “A Study on the Influence of Local Buckling on the Strength of Structural Core Sandwich Structures", Multiaxial Fatique and Deformation Testing Techniques, ASTM STP, Vol. 1280, ed. S. Kalluri and P. J. Bonacuse, pp. 264-289, 1997. 15. A. N. Palazotto and S. T. Dennis, "Nonlinear Analysis Analysis of Shell Structures", Structures", AIAA Education Series, pp. 116 - 154, 1992. 16. R. D. Cock, D. S. Malkus Malkus and M. E. Plesha, "Concepts "Concepts and Applications of Finite Element Analysis", John Wiley & Sons, 1989. 17. B. S. Westerlind and L. A. Carlsson, "Compressive Response of Corrugated Board", TAPPI, 75 (7), pp. 145-154, 1992.
18
Appendix
Table A1. Specifications, test test results and FEM-estimates FEM-estimates for handmade B-flute boxes, COV was about 5 % for all tested boxes. Box no.
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51
Length
Width
Height
(mm)
(mm)
(mm)
200 300 350 400 400 400 400 400 450 500 535 550 300 300 300 300 300 400 500 600 600 300 300 300 300 300 400 400 400 375 420 450 480 510 540 540 540 570 590 300 300 300 300 300 400 500 600 600 400 500 600
150 225 350 300 300 300 300 300 250 200 400 150 300 300 300 300 300 400 300 200 200 300 300 300 300 300 400 400 400 250 280 300 320 340 360 360 360 380 390 300 300 300 300 300 400 300 200 200 400 300 200
300 300 300 300 200 400 600 800 300 300 300 300 100 200 300 450 600 400 400 400 300 100 200 300 450 600 400 600 133 400 400 400 400 400 200 400 600 400 400 100 200 300 450 600 400 400 400 300 400 400 400
Flute
Board grade
B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B B
no. of samples
125MK/150RF/125T 125MK/150RF/125T 125MK/150RF/125T 125MK/150RF/125T 125MK/150RF/125T 125MK/150RF/125T 125MK/150RF/125T 125MK/150RF/125T 125MK/150RF/125T 125MK/150RF/125T 125MK/150RF/125T 125MK/150RF/125T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 280K/150RF/280K 280K/150RF/280K 280K/150RF/280K Total
19
6 6 6 6 6 6 6 6 6 6 6 6 10 10 10 10 10 10 10 10 10 20 20 20 20 19 20 17 20 10 10 10 10 10 9 10 10 10 10 10 10 10 10 10 10 10 10 10 19 20 20 566
Mean BCT
FEM
(N)
(N)
1801 1939 2072 2020 1924 1820 1839 2042 1842 1646 1931 1631 2216 1794 1981 1657 1657 2148 2040 1893 1863 2703 2702 2567 2301 2370 2597 2377 2754 3020 2471 2618 2785 3118 2726 2540 2824 3158 2805 2716 2657 2795 2569 2481 3109 3442 2932 2893 4762 4866 3987
1794 1908 1942 1881 1824 1892 1915 1950. 1850. 1723 1964 1554 2543 1866 1954 1959 1975 1953 1903 1771 1815 3229 2493 2629 2608 2656 2606 2642 2675 2670 2595 2608 2635 2710 2697 2743 2782 2740 2740 2839 2425 2694 2625 2722 2887 2479 2251 2234 4240 3814 3378 Mean CV %
BCT/FEM
1.004 1.015 1.066 1.074 1.054 0.962 0.960 1.047 0.996 0.955 0.983 1.049 0.871 0.962 1.013 0.846 0.839 1.099 1.072 1.069 1.026 0.837 1.084 0.976 0.882 0.892 0.997 0.899 1.029 1.131 0.952 1.003 1.057 1.151 1.011 0.926 1.015 1.152 1.024 0.956 1.096 1.037 0.979 0.911 1.077 1.388 1.302 1.295 1.123 1.276 1.180 1.0314 11.4
Appendix
Table A2. Specifications, test test results and FEM-estimates FEM-estimates for handmade C-flute boxes, COV was about 5 % for all tested boxes. Box no.
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49
Length
Width
Height
(mm)
(mm)
(mm)
300 300 300 300 300 300 300 300 400 500 600 600 300 300 300 300 300 400 400 400 400 500 600 375 395 420 450 480 510 540 540 540 570 590 300 300 300 300 300 300 360 400 400 500 600 600 400 500 600
300 300 300 300 300 300 300 300 400 300 200 200 300 300 300 300 300 400 400 400 400 300 200 250 260 280 300 320 340 360 360 360 380 390 300 300 300 300 300 300 240 200 400 300 200 200 400 300 200
100 200 300 450 600 150 300 450 400 400 400 300 100 200 300 450 600 400 600 133 400 400 400 400 400 400 400 400 400 200 600 400 400 400 300 100 200 300 450 600 300 300 400 400 400 300 400 400 400
Flute
Board grade
no. of samples
Mean BCT (N)
(N)
C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C C
140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 140T/112RF/140T 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 180L/112RF/180L 186K/112RF/180T 186K/112RF/180T 186K/112RF/180T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/112RF/200T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 186K/150RF/180T 280K/150RF/280K 280K/150RF/280K 280K/150RF/280K
10 10 10 10 10 30 30 30 10 10 10 10 20 19 20 20 20 20 19 20 19 20 19 10 10 10 9 10 10 10 10 20 10 10 29 10 10 10 10 10 30 29 10 10 10 10 19 20 18
2932 2177 2187 2138 2118 2546 2411 2246 2540 2442 2324 2226 3290 2996 3168 2806 2889 3417 3178 4134 3488 3289 2929 3275 3265 3403 3805 3913 3971 4638 3775 4011 4226 4324 3606 4423 3707 3373 3285 3167 3730 3545 3864 4030 3462 3275 6596 6315 5811
4931 2597 2873 2857 2947 2932 2873 2857 2812 2685 2368 2385 6113 3481 3844 3807 3928 3738 3720 4762 3701 3369 2906 3775 3750 3746 3720 3735 2764 3197 3904 3807 3835 3841 4170 4959 3542 4170 3818 3904 3393 3026 3846 3491 3002 2945 6097 5306 4465
Total
20
750
FEM
Mean CV %
BCT/FEM
0.595 0.838 0.761 0.748 0.719 0.868 0.839 0.786 0.903 0.909 0.981 0.933 0.538 0.861 0.824 0.737 0.735 0.914 0.854 0.868 0.942 0.976 1.008 0.867 0.871 0.908 1.023 1.048 1.437 1.451 0.967 1.053 1.102 1.126 0.865 0.892 1.046 0.809 0.860 0.811 1.099 1.172 1.005 1.155 1.153 1.112 1.082 1.190 1.301 0.950 19.6
Appendix
Table A3. Material data for liners and flutings. 125 Testliner Grammage (g/m 2 ) Thickness ( µ m) T ensile stiffness MD (M N/m) T ensile stiffness CD (M N/m) Tensile strength M D (kN /m) Tensile strength CD (kN/m) SCT MD (kN/m) SCT CD (kN/m) 125 Kraftliner Grammage (g/m 2 ) Thickness ( µ m) T ensile stiffness MD (M N/m) T ensile stiffness CD (M N/m) Tensile strength M D (kN /m) Tensile strength CD (kN/m) SCT MD (kN/m) SCT CD (kN/m) 140 Testliner Grammage (g/m 2 ) Thickness ( µ m) T ensile stiffness MD (M N/m) T ensile stiffness CD (M N/m) Tensile strength MD (kN/m) Tensile strength CD (kN/m) SCT MD (kN/m) SCT CD (kN/m) 140 Kraftliner Grammage (g/m 2 ) Thickness ( µ m) T ensile stiffness MD (M N/m) T ensile stiffness CD (M N/m) Tensile strength M D (kN /m) Tensile strength CD (kN/m) SCT MD (kN/m) SCT CD (kN/m) 180 Testliner (L) Grammage (g/m 2 ) Thickness ( µ m) T ensile stiffness MD (M N/m) T ensile stiffness CD (M N/m) Tensile strength M D (kN /m) Tensile strength CD (kN/m) SCT MD (kN/m) SCT CD (kN/m)
121 20 0 0.968 0.363 7.744 2.783 3.509 1.815
186 Kraftliner Grammage (g/m 2 ) Thickness ( µ m) T ensile stiffness MD (M N/m ) T ensile stiffness CD (M N/m) Tensile strength M D (kN/m) Tensile strength CD (kN/m) SCT MD (kN/m) SCT CD (kN/m)
182 22 3 1.6 81 0.6 14 16.59 7.05 6.07 3.51
125 16 4 1.374 0.503 12.8 4.15 4.74 2.53
200 Testliner Grammage (g/m 2 ) Thickness ( µ m) T ensile stiffness MD (M N/m ) T ensile stiffness CD (M N/m) Tensile strength M D (kN/m) Tensile strength CD (kN/m) SCT MD (kN/m) SCT CD (kN/m)
200 31 5 1.5 37 0.5 01 12.34 4.24 5.68 3.16
136 22 5 0.83 0.46 6.4 3.8 3.33 2.38
280 Kraftliner Grammage (g/m 2 ) Thickness ( µ m) T ensile stiffness MD (M N/m ) T ensile stiffness CD (M N/m) Tensile strength M D (kN/m) Tensile strength CD (kN/m) SCT MD (kN/m) SCT CD (kN/m)
280 33 8 2.3 76 0.8 75 23.8 9.76 8.65 5.06
139 19 5 1.39 0.49 14.0 5.3 5.1 2.7
112 Recycled fibre fluting Grammage (g/m 2 ) Thickness ( µ m) T ensile stiffness MD (M N/m ) T ensile stiffness CD (M N/m) Tensile strength M D (kN/m) Tensile strength CD (kN/m) SCT MD (kN/m) SCT CD (kN/m)
114 19 9 0.9 11 0.3 54 7.32 2.9 3.8 2.12
178 28 3 1.19 0.55 9.97 4.63 4.98 3.29
150 Recycled fibre fluting Grammage (g/m 2 ) Thickness ( µ m) T ensile stiffness MD (M N/m ) T ensile stiffness CD (M N/m) Tensile strength M D (kN/m) Tensile strength CD (kN/m) SCT MD (kN/m) SCT CD (kN/m)
151 23 1 1.1 42 0.4 49 8.95 3.45 4.82 2.76
21
